Technology Review 2012 MICA(P) 039/02/2012

Technology Review 2012 MICA(P) 039/02/2012

About KOMtech Launched in December 2007, Keppel Offshore & Marine Technology Centre (KOMtech) is an extension and strengthening of Keppel Offshore & Marine’s (Keppel O&M) research and development initiatives. Its mission involves promoting innovation, establishing technology foresight in alternative energy applications as well as developing designs systems and critical equipment for rig and ships.

KOMtech complements and augments the work of three design and engineering units within Keppel O&M - Offshore Technology Development, Deepwater Technology Group and Marine Technology Development. Leveraging existing and proprietary technologies, and collaborating with universities, research institutes and industry partners worldwide, KOMtech continues to develop innovative solutions that are commercially viable and adaptable to the needs of the industry.

Ongoing research efforts at KOMtech span topics such as Arctic structures, drilling systems and mini-LNG supply-chain for associated gas. Current areas of research also include the development of a slim drillship as well as offshore wind energy - related solutions including self-installing platforms for substations, wind turbine foundations, wind turbine installation vessels, marine emission control and cable laying vessels.

Chairman’s Message

Over the years, we have harnessed innovation and technology to address the varied and changing demands of drilling companies and operators worldwide. This has enabled Keppel Offshore & Marine (Keppel O&M) to progress up the value chain, keep our competitive edge and sustain leadership. We continue our focus on developing products that are commercially viable and relevant to the market’s needs.

These serve as fundamental principles for KOMtech in the research and development (R&D) of new products as well as innovation and enhancement of existing products and processes. With a team of about 65 researchers, KOMtech augments existing design and engineering activities within Keppel O&M by providing new specialised capabilities and upstream R&D of new product designs.

Technology Review 2012 brings KOMtech’s multidisciplinary R&D expertise to bear, with a closer look at the trends of jacket substructures for offshore wind turbines, hose transfer handling system for Liquefied Natural Gas offloading and the wet scrubbing process for marine emission control. Some of the products developed are new inventions while others are alternative concepts that address technical challenges in actual projects.

Prior issues of the Review have also shored up substantial interest from the offshore and marine industry, which includes our customers, professional institutions and academia. Several of KOMtech’s papers have also been presented at the European Wind Energy Association 2011 Conference, Offshore Support Vessel 2011 Conference and Jackup 2011 conference as well as published by the Society of Naval Architects and Marine Engineers Singapore.

My appreciation goes to KOMtech’s team of able researchers, led by Dr Foo Kok Seng Executive Director (Shallow Water Technology) and Director/Advisor, Mr Charles Foo. I would also like to take this opportunity to welcome Mr Aziz Amirali Merchant, Executive Director (Deepwater Technology) onboard. The publication brings together inputs from our research partners and customers, who represent some of the best minds and authorities in our industry. It is my hope that this publication will help synergise further ideas and innovations for the industry.

I trust that you will find this fourth edition of the publication thought-provoking and stimulating.

Yours truly,

Choo Chiau Beng Chairman Keppel Offshore & Marine

Contents

About KoMtech

1 ChAirMAn’s MessAge

5 editoriAl note

7 CorPorAte highlights

reseArCh highlights

11 Hose Transfer Handling System for Offshore LNG Offloading A transfer method is used to handle flexible composite cryogenic hoses for LNG offloading in offshore environment

19 Development of OTD Hydraulic Jacking System An improved design of the hydraulic jacking system for jackup vessels in offshore installation market

25 Simulating LNG Liquefaction Plant Start-up This paper investigates the dynamic simulation of an LNG liquefaction plant to better understand its operational behavior and control of the process

31 Ballast Water Treatment System Selection A model to assist ship owners in selecting an appropriate ballast water treatment system for their vessels

39 Wet Scrubbing Process for Marine Emission Control KOMtech has developed a high performance effective wet scrubbing process for SOx removal from marine emission

49 CFD Simulation of Water Oscillations in the Moonpool of a Drillship Advanced computational fluid dynamics simulation approach with high performance computing techniques are applied to simulate water oscillations in the moonpool of a drillship

About KoMtech

1 ChAirMAn’s MessAge

5 editoriAl note

7 CorPorAte highlights

reseArCh highlights

11 Hose Transfer Handling System for Offshore LNG Offloading A transfer method is used to handle flexible composite cryogenic hoses for LNG offloading in offshore environment

19 Development of OTD Hydraulic Jacking System An improved design of the hydraulic jacking system for jackup vessels in offshore installation market

25 Simulating LNG Liquefaction Plant Start-up This paper investigates the dynamic simulation of an LNG liquefaction plant to better understand its operational behavior and control of the process

31 Ballast Water Treatment System Selection A model to assist ship owners in selecting an appropriate ballast water treatment system for their vessels

39 Wet Scrubbing Process for Marine Emission Control KOMtech has developed a high performance effective wet scrubbing process for SOx removal from marine emission

49 CFD Simulation of Water Oscillations in the Moonpool of a Drillship Advanced computational fluid dynamics simulation approach with high performance computing techniques are applied to simulate water oscillations in the moonpool of a drillship

63 New Painting Technologies for Productivity Improvement Discusses how technology is employed in the production activities of the yards to augment productivity of the painting process

67 Fatigue Assessment of Three-Chord Jacket Substructure for Offshore Wind Turbines in the North Sea A study is carried out on the fatigue assessment of a three-chord jacket substructure to determine if simplified assumptions used in Damage Equivalent Load method will lead to overdesign

FeAtured ArtiCles

79 Comparison of SNAME T&RB 5-5A Framework and a Plasticity-based Spudcan Model for Jackup Foundation Assessments A comparison is done using a plasticity model with the SNAME method for jackup foundation assessments

93 Use of Field Penetrometer Data for an Integrated Jackup Installation System An integrated jackup installation system is used to assist operators in deciding how to prevent or mitigate potential geotechnical hazards

113 Installing Offshore Wind Turbines in Harsh Environments Outlines how wind turbine blades are handled on a large four-legged jackup vessel and installed offshore safely and efficiently

125 New Generation Deep Water OSVs for Oil & Gas Operation The paper gives an insight into the features of new generation PSVs and AHTS for supporting more complex deepwater field developments and describes two new proprietary designs from MTD

131 Emulsified Fuel System for Marine Diesel Engines The paper presents emulsified fuel system as a compelling solution to ship owners for reducing fuel consumption

4 KOMtech Technology Review 2012

Editorial Note

Editorial Note 5

It gives me great pleasure to introduce Technology Review, an annual publication produced by Keppel Offshore & Marine Technology Centre (KOMtech) to share its extensive research work.

This 2012 edition highlights KOMtech’s ongoing and new research areas, including its solutions for hose transfer handling system for offshore LNG offloading as well as marine emissions control.

In addition, Technology Review 2012 delves into innovations and technologies for harnessing offshore wind energy. More specifically, it includes papers on KOMtech’s hydraulic jacking system for wind installation vessels as well as fatigue assessments for jacket substructures for offshore wind turbines.

Two new chapters have been added to this 2012 edition to help readers better understand KOMtech and its work. “Featured Articles” comprise research papers published or presented at conferences and seminars by KOMtech’s researchers. This section also includes select papers from Keppel Offshore & Marine’s (Keppel O&M) various business units and its industry partners. Meanwhile, “Corporate Highlights” give readers a glimpse of events which have taken place at KOMtech over the past year.

Through Technology Review, KOMtech hopes to stimulate further discussions and innovations as well as showcase its capabilities and solutions. Every attempt has been made to ensure accurate and factual representation of the research presented. It is important to note that most of these research projects are still ongoing, and opinions and views expressed in the papers are that of the authors and are not necessarily the official position of KOMtech.

We deeply appreciate the contributions from industry partners, universities and Keppel O&M’s business units towards our research work and the development of this publication. The authors and editors of Technology Review 2012 have put in many hours to ensure that this publication is insightful and thought-provoking. We would like to thank everyone involved for their hard work and support.

Wayne Yap Chairman, Editorial Committee

editoriAl CoMMittee

Advisors Mr Charles Foo and Dr Foo Kok Seng

Editorial members Mr Wayne Yap, Dr Henry Krisdani, Mr Alex Tan, Dr Bernard How,

Dr Prapisala Thepsithar, Dr Huynh Le Ngoc Thanh and

Ms Wong Chai Yueh (Keppel Group Corporate Communications)

6 KOMtech Technology Review 2012

Corporate Highlights 7

Corporate Highlights

Top students from six German universities visited KOMtech on 31 March 2011. Majoring in the fields of naval architecture, marine engineering and offshore engineering, the students are members of the German Society for Maritime Technology.

Mr Charles Foo, Director/Advisor of KOMtech, hosted the delegation. Through a corporate presentation and a tour of KOMtech’s facilities as well as Keppel Offshore & Marine’s (Keppel O&M) yards in Singapore, the students gleaned insights to Keppel O&M’s businesses and research activities.

Members of the german society for Maritime technology visited KoMtech on 31 March 2011

KoMtech hosted Msc students from delft university of technology

Seeking a deeper understanding of the offshore sector and career opportunities within the industry, 20 MSc Offshore Engineering students from Delft University of Technology in Netherlands visited KOMtech. This visit formed part of their overseas study tour.

Dr Foo Kok Seng, Executive Director (Shallow Water Technology), outlined KOMtech’s technology focus and research findings during the visit. He also shared his perspectives on the future of the offshore and marine industry and highlighted career opportunities within Keppel O&M. The visit concluded with a tour of Keppel FELS.

International exchange

8 KOMtech Technology Review 2012 CORPORATE HIGHLIGHTS

KOMtech’s Offshore Structure Group Assistant Manager, Dr Michael Perry, presented at the 30th International Conference on Ocean, Offshore and Arctic Engineering (OMAE) held in Rotterdam, the Netherlands, in June 2011. He was also asked to speak at the Outreach for Engineers Specialty Forum, which is part of the conference.

The specialty forum aims to provide a broad overview of the industry as well as highlight potential career opportunities within the field. In his presentation, Dr Perry shared on KOMtech’s research areas as well as his passion for his work.

KoMtech’s dr Michael Perry inspired young minds to take on the challenges of the offshore and marine industry

A workshop on regulatory outlook & proposed response for Mobile offshore drilling units

KOMtech organised a workshop on Regulatory Outlook and Proposed Responses for Mobile Offshore Drilling Units (MODUs) for Keppel O&M’s business units.

Mr AK Seah, American Bureau of Shipping (ABS), Vice President for Environmental Solutions, spoke at the workshop. The session focused on marine regulations coming into force in the near future as well as their impact on rig designs.

More specifically, the workshop provided more details on the IMO MODU Code 2009, ballast water management regulations, energy efficiency regulations, ship recycling conventions as well as maritime labour conventions.

KOMtech’s Corrosion and Material Engineer, Mr Xu Da Qin, shared his knowledge on basic corrosion and seawater cooling systems corrosion with Keppel O&M’s various business units including Keppel FELS, Keppel Shipyard, Keppel Singmarine and Keppel Smit Towage.

Held in September 2011, the sharing session touched on the basics, causes and prevention methods of corrosion on common engineering materials as well as corrosion prevention measures for seawater heat exchanger tubes and pipes.

Corrosion sharing session between KoMtech and Keppel o&M’s business units

Knowledge sharing

KoMtech participates in ntu’s JiP in offshore renewables

KOMtech will represent Keppel O&M in Nanyang Technology University’s (NTU) Joint Industry Programme in Offshore Renewables, a collaboration between academia and industry to develop more efficient and effective offshore wind and marine power generation systems.

Besides Keppel O&M, other multinational companies that have signed up for this programme include Rolls Royce, DNV and Vestas.

The Joint Industry Programme, hosted at Energy Research Institute @ NTU (ERI@N) will have over 25 industry and application-oriented research projects over the next three years, grooming 35 PhD students, Masters students, and research associates.

Corporate Highlights 9

Partnership with academia

Taking in interns from various universities and tertiary institutes, KOMtech helps to cultivate aspiring researchers. With ample opportunities to learn from KOMtech’s researchers, these students expand their knowledge and gain in valuable working experience.

A former intern at KOMtech, Nicholas Yap, who is presently pursuing a Bachelor of Engineering (Chemical) degree from the University of New South Wales, said, “My internship experience with KOMtech provided me with opportunities to complement theoretical understanding with practical know-how. In addition, I gained knowledge which can be applied in courses I will take in the near future. The exposure to the office setting and the field site prepare me for a smooth transition to the workplace.”

Grooming Talent

10 KOMtech Technology Review 2012 ReseaRch highlights

No part of the materials published in this journal may be reproduced, stored in a retrieval system or transmitted in any form whatsoever without the prior written permission of KOMtech

Hose Transfer Handling System for Offshore LNG Offloading 11

the tRansfeR of natuRal gas in liquid phase, in the foRm of liquefied natuRal gas (lng), between a LNG carrier and a floating LNG facility offshore at cryogenic temperatures of about -162˚C is complex and challenging. The motions at each end of the transfer chain demand a flexible solution so as to transfer the cryogenic hydrocarbon fluid safely.

KOMtech has developed a transfer method to handle flexible composite cryogenic hoses for LNG offloading between a LNG carrier and a floating LNG facility. This paper outlines the novel key enabling technologies involved, which provide a cost effective solution in the transfer of LNG in an offshore environment.

Hose Transfer Handling System for Offshore LNG Offloading

alex tan Kah Keong, B.Eng

foo Kok seng, PhD, B.Eng

asbjorn moRtensen, Dr.Ing, M.Eng

Ruud J.p. haneVeeR, B.Sc


INTRODUCTION For many decades, natural gas was just a by-product of oil production. The reason was associated with the transportation of the gas especially offshore where long gas pipelines are not feasible economically. Liquefaction of natural gas allows for mobility in terms of transporting the Liquefied Natural Gas (LNG) anywhere in the world with LNG ships or LNG carriers (LNGC). LNG, a natural gas in liquid phase at cryogenic temperatures approximately -162˚C, has developed into a worldwide commodity[1] and is the fastest growing hydrocarbon fuel. The development of LNG has been spurred by the massive amount of stranded gas, environmental regulations that ban flaring, the development for alternative energy mix and price spikes in natural gas prices.

LNG transfer systems are needed as isolated energy markets cannot be connected to pipeline infrastructures whilst international legislation favours the use of LNG as clean fuel for shipping.

KOMtech’s LNG offloading solution is developed to cater for a wide range of applications in the LNG market with the primary objectives to cope with the increased safety requirements for medium to small scale storage, bunkering and distribution facilities, and to enable the offshore production of stranded/associated gas at remote locations without pipeline grids nearby.

OFFSHORE LNG TRANSFERThere are several factors that resulted in the transfer of LNG from near shore/onshore terminal to offshore. The reasons are safety concerns to the community near LNG processing facilities, lack of land area in congested port, limitation for capacities increase in future, congested shipping at the port, security risks and concerns with regards to terrorism, CAPEX and OPEX considerations, and insufficient deepwater close to the coast at small isolated markets where LNGCs are unable

to berth. These resulted in offshore LNG transfer, process and storage establishing itself as an attractive proposition. Furthermore, there are renewed efforts being initiated to develop offshore stranded gas reserves with floating LNG facilities.

The trend for production and storage of LNG is moving further offshore and hence, it is necessary that the transfer system for LNG to be safe, efficient, reliable and easy to operate in an offshore environment.

CURRENT TECHNOLOGY AND APPLICATIONA typical storage capacity range for a small to medium sized LNGC is 8,000 m3 - 60,000 m3, while a large LNGC can carry approximately 100,000 m3 to 300,000 m3 or more LNG.

Near shore/onshore terminalLNG transfer utilising robotic rigid pipes assembly, more commonly known as loading arms, are the common choices of transfer system used in near shore/onshore terminals. The loading arm comes in 8”, 10” or 16” diameter pipes. The rigid pipes are hinged together with swivels, allowing a six-degree-of-freedom and a counterweight system, is commonly incorporated. This allows the loading arm to weathervane according to the environmental condition. A large LNGC traditionally employs 4 - 6 loading arms with 16" flanges for an offloading operation at the terminal.

Ship-to-ship In benign offshore location, ship-to-ship transfer of LNG has been performed with composite cryogenic hoses. In the current industry practice, the 16" manifolds of the ships are each connected by two composite cryogenic hoses of 8" diameter between them. A typical transfer operation utilises three pairs of manifolds connection with six composite hoses between the ships, which achieves a LNG offloading rate of approximately

The trend for production and storage of LNG is moving further offshore and hence, it is necessary that the transfer system for LNG is safe, efficient, reliable and easy to operate in an offshore environment.

6000 m3/h. The two ships are typically in a side-by-side mooring configuration with floating pneumatic fenders keeping a minimum separation distance of 3 m to 5 m between them.

Offshore environmentThe transfer of LNG offshore is either between a LNGC and a fixed LNG facility or between a LNGC and a floating LNG facility. The LNG facility can be a production (liquefaction), processing, re-gasification or storage (import/export) facility. The safe and reliable transfer of LNG offshore in an open sea with possibly severe environment has never been done although there are a few transfer methods being developed and conceptualized in the industry[2]. The offshore LNG transfer configuration may be in a side-by-side mooring or a tandem arrangement. These transfer methods are mainly developed and conceptualized based on the use of aerial hanging or floating LNG hoses although the method of using a boom structure with robotic rigid pipes has also been developed for tandem LNG offloading.

CHALLENGES IN OFFSHORE LNG TRANSFERThere are several challenges related to LNG transfer offshore. The main challenges associated with LNG transfer in harsh environmental conditions with significant wave heights of up to 5.5 m with duration between 8 and 12 seconds, as well as higher significant wind and current loads for LNG offloading duration of 18 to 24 hours[3]. There are limitations in the current LNG transfer technology that is being applied in benign environmental conditions, therefore it is not able to be adapted directly to offshore LNG transfer which involves high dynamic motions. Operational issue related to the high dynamic motions involved in offshore LNG transfer is also an important safety concern to be further evaluated.

The use of traditional loading arms for offshore LNG transfer is often limited to benign sea conditions. The traditional loading arms are unable to handle the high dynamic motions and higher wind speed in the rough sea due to the high bearing and bending stresses associated on the rigid pipes and swivels. In addition, the investment and maintenance cost (the lifecycle cost) of the loading arms are considered too high for small to medium scale LNG transfer applications.

Current handling of composite cryogenic hose for ship-to-ship LNG transfer in benign offshore environment utilizes the cargo crane available on the LNGC for hoisting a hose at a time to the adjacent LNGC. Winches might also be required to assist with the connections of the hoses between two vessels. This operation requires a significant number of crew members and high amount of manual labour. The downtime associated with such operations is also high. In addition, performing such a transfer operation at harsher offshore environment poses safety hazards.

A spot market for LNG is developing according to the trends in the current market. Therefore, in order to allow flexible and efficient operation of LNG facilities, it is important for a common connector technology[4] to be available, so that LNGCs of opportunity can be handled[5]. Therefore, to uphold the suitability of non-dedicated vessels at the LNG facilities offshore, the transfer of LNG should take place at the midship manifold with minimum modification on the LNGC for adaption of the transfer system. Other considerations include the compactness of the transfer system design for offshore application and maintenance/inspection required for high operational uptime.

The commercial solution for offshore LNG transfer should be developed to take advantage of the specific location where the transfer system is to be applied whilst addressing the main challenges of operating in such an environment. There are several variations of LNG transfer system developed for both side-by-side and tandem LNG offloading with the point taken that tandem solutions being unsuitable for non-dedicated vessels. The technology is available but significant challenges remain to ensure safe operations and high uptime as it is essential to provide customers with a reliable service at the LNG facilities as well as to extend the safe limits for LNG offloading in rough seas.

KOMtech LNG OFFLOADING SOLUTIONKOMtech LNG hose transfer offloading system is designed with handling aerial composite cryogenic hose for offshore LNG transfer. The choice of employing composite cryogenic hose in an aerial configuration for offshore LNG



transfer boils down to several reasons. Multi layers composite hoses are known to be very flexible compared to stainless steel braided hoses. Due to the low temperatures at approximately -162˚C when handling LNG, such flexible composite hoses have only just recently become technically and commercially viable for LNG application. They have been tested for their reliability, which they excelled. These flexible composite cryogenic hoses are now commercially available through a few dedicated hose manufacturers. In comparison, floating hoses technology for LNG transfer has yet to be widely available in the industry. There are also questions about the commercial viability for small to medium scale LNG transfer applications.

Aerial hoses are also believed to be a better alternative at offshore compared to loading arms as there are advantages such as reduced cost and improved flexibility that allows LNG transfer in the rough open sea. Furthermore, KOMtech’s solution offers a wider allowable operating envelope (Figure 1) especially in the relative vessel surge direction compared to loading arms which have limitations in weathervane along the transverse direction. This is an important consideration for LNG transfer offshore.

KOMtech LNG hose transfer offloading system targets the small-medium scale LNG services. It is also developed to overcome the challenges in LNG transfer offshore, particularly in the area of safety,

Figure 1. Allowable operating envelope (Profile View & top View) of KoMtech’s solution

KOMtech’s LNG hose transfer offloading system targets the small-medium scale LNG services. It is also developed to overcome the challenges in offshore LNG transfer, particularly in the area of safety, operational ease, adaptability and compactness of the system for offshore application.

2m relative heave up

2m allowable relative freeboard

2m relative heave down

typical ship-to-ship separation

lngc lng facility

operational ease, adaptability and compactness of the system for offshore application. The system utilizes the standard 8” composite cryogenic hose technology available in the market. Such medium sized hoses enable flow transfer rates of around 2000 m3/h with 2 hoses, as opposed to 6000 m3/h or more with 4 - 6 traditional 16” LNG loading arms typically employed for large LNGC. For small - medium sized LNGC shuttling between many remote locations with smaller receiving tanks, the lower cost system associated with aerial hoses is a more suitable solution as a mini LNGC will not engage the LNG loading arms for long and the small to medium cargo volume does not justify the higher cost of loading arm technology. In addition, the aerial hose solution provides improved operational window which means less downtime during bad weather conditions.

KOMtech LNG HOSE TRANSFER OFFLOADING SYSTEM The target application of KOMtech hose transfer system is on LNG transfer from a small - medium sized LNGC to a floating LNG facility with a side-by-side mooring configuration utilizing the midship manifolds of the LNGC. The system can also be configured for LNG transfer to a bow loading LNGC in a tandem arrangement. In both instances, the same proposed LNG hose transfer technology can be used for the offloading operation, although actual vessel separation distance and hose length may vary according to the location and its environmental condition.

The hose transfer offloading system can be divided into three main modules (Figure 2); Lifting Appliance, Transfer Skid and Pipe Deck. Each

Figure 2. KoMtech lng hose transfer offloading system


lifting appliance

pipe deck

transfer skid


module serves a particular function to enable the complete system to perform to the required design specifications.

The lifting appliance is the mechanical handling machine of the system. It will typically be installed on the floating LNG facility. The primary function of the lifting appliance is to hoist the transfer skid (including the attached composite cryogenic hoses) to the adjacent LNGC for connection. The lifting appliance will also provide constant tension targeting guide lines to assist in the positioning of the transfer skid on the adjacent LNGC.

The transfer skid houses multiple hoses and their components together in one module thereby allowing multiple hoses to be transferred simultaneously in a single lifting operation by the coupled lifting appliance of the system. The standard configuration of the transfer skid consists of three composite cryogenic hoses, two hoses for LNG transfer and one hose reserved for vapour return. Other configurations with more or less hoses are possible. Each composite hose is coupled to an Emergency Release Coupling (ERC) which is assembled with an isolation valve and a Quick Connect/Disconnect Coupling (QC/DC) flange in a hard pipe assembly on the transfer skid.

The pipe deck will be installed at the midship section of the LNGC. It consists of pipe spools for connection between the LNGC midship manifolds and the QC/DCs on the transfer skid. In addition, the pipe deck consists of two guide posts for stabbing into the guide funnels of the transfer skid with the assistance of the guide lines from the lifting appliance. This is to target the transfer skid to the right position on the LNGC for the engagement of the QC/DCs and pipe spool flanges for LNG transfer.

The installation of the pipe deck on the LNGC is considered to be a minimum modification on the vessel for the adaption to this transfer technology. The pipe deck can be easily removed at any point in time from the LNGC grating deck if a different transfer system needs to be employed at another LNG terminal.

A Linked Emergency Shut Down (ESD) and Emergency Release System (ERS) are incorporated into the transfer system which is used to execute a controlled rapid shut down and disconnection of the LNG transfer operation in emergency situations. Emergency alarms will trigger if the transfer operation exceed the prescribed allowable working envelope of the system. The alarm 1st step detection triggers the ESD sequence, among other things, transfer pumps are stopped and ESD valves are shut, terminating any LNG transfer. The alarm 2nd step may be detected thereafter which triggers the ERS sequence whereby the ERCs valves are closed and break away of the ERCs initiated. The ESD & ERS initiation can be activated by hardwires which connect to both the vessels. In the event whereby the two vessels lose positioning and inadvertently drift apart beyond the allowable working envelope during LNG transfer, the cables are pulled to trigger alarm 1st step and 2nd step at different intervals.

BENEFITSKOMtech hose transfer offloading system is developed to bridge the gaps in the LNG market today. The benefits of this development are listed as follows:

• Versatilityasthetransfersystemisabletobe applied at terminal, near shore or offshore.

• Flexibilityofcompositecryogenichoses allow for high environmental criteria for LNG transfer operation in a dynamic motion environment such as in the rough open sea.

• Potentiallylessdowntimeforworseweather condition or monsoon period when LNG transfer operation is not possible with conventional methods. This results in improved operational window.

• CostefficientsolutionwithreducedCAPEX for small-medium scale LNG application at remote locations.

• ReducedOPEXwithmechanizedoperation requiring minimal operators.

• Enhancedsafetybyminimizinghuman intervention and assistance during the connecting and disconnecting operations especially during harsh weather conditions.

• Reliablesystemconsistingofproven cryogenic components.

• Reducedoperationaltimewithefficient transport of multiple hoses at any one time simultaneously.

• Efficientandfastconnectioncanbeachieved with guiding/targeting system and quick connecting couplers.

• Simpleandeasytooperate.

• Adaptabilityofthesystemwithminimum modifications to non-dedicated vessels.

• Applicableforsidebysideortandemvessel arrangement.

• Compactnessofthesystemsavingdeckspace for application offshore.

• Emergencysolutionintermsofsafe disconnection and minimum possibility of spillage on vessel deck implemented.

• Numberofhoseconfigurationcustomizable.

• Wideallowableoperatingenvelopewhichis adaptable according to location requirement.

• Thetransfersystemisdevelopedaccording to the industry codes and standards as well as applicable rules for LNG transfer.

CONCLUSIONAs the gas market expands and development of stranded gas market picks up pace in the years ahead, the demand for a safe, reliable and cost efficient LNG transfer system offshore will increase. The development of KOMtech LNG hose transfer offloading system aims to overcome the challenges faced by operators today. The patent pending design has also obtained Approval-In-Principle certification for conceptual engineering, operating envelope, loading computation and recommended operating procedure from classification society, Lloyd’s Register.

It should be noted that a major factor to the successful development of a suitable offshore LNG transfer system lies in the choice of the mooring configuration. There are several recommendations on the minimum separation between vessels based on the sea state. The primary focus of KOMtech LNG hose transfer offloading system in the future will be on the manipulability of the system to higher sea states as gas development moves further and deeper offshore.

AuTHOR’S CONTACT [email protected] • [email protected]

REFERENCES [1] L. Poidevin, 2 June 2010, Addressing the Challenges of Offshore LNG Transfer, http://www.epmag.com/Production-Field-Development/Addressing- challenges-offshore-LNG-transfer_60862, Accessed 20 December 2011.

[2] L. Poldervaart, J. Ellis, The Transfer of LNG in Offshore Conditions. Same Song – New Sound, 15th International Conference & Exhibition on Liquefied Natural Gas (LNG 15), Barcelona, Spain, April 2007.

[3] S. Hoog, H. Koch, R. Huhn, C. Frohne, J. Homann, G. F. Clauss, F. Sprenger, D. Testa, LNG Transfer in Harsh Environments – Introduction of a New Tandem Mooring Concept, AIChE Spring National Meeting, Tampa, Florida, April 2009.

[4] D. McDonald, C.H. Chiu, D. Adkins, Comprehensive Evaluations of LNG Transfer Technology for Offshore LNG Development, LNG-14, Doha, Qatar, March 2004.

[5] J. de Baan, M.H. Krekel, R. Leeuwenburgh, M.M. McCall, Offshore Transfer, Re-Gasification and Salt Dome Storage of LNG, Offshore Technology Conference, Houston, Texas, u.S.A, May 2003.




Development of OTD Hydraulic Jacking System 19

theRe is a need in the offshoRe installation maRKet foR JacKup Vessels with stRong JacKing systems, featuring long lifespan, high jacking speed and continuous reliable operation with minimum down time. Existing hydraulic jacking system designs face problems with jacking speed, limited space for jack cases on main deck and complexity in pin engagement operations. This paper focuses on the design of a hydraulic continuous jacking system, using a shorter load path for transmitting the jacking forces to reduce the size of the jack case. The pin engagement sequence utilises a minimum number of intermissions, therefore simplifying the electrical control system and providing smooth jacking. Enhanced reliability for the vessel is achieved by employing an additional emergency locking unit to fix the hull to the legs in case of a cylinder failure.

Development of OTD Hydraulic Jacking System

michael Kyaw nyunt, B.Eng

shi degang, B.Eng Offshore Technology Development

strahil mihayloV, M.Eng, B.Eng Offshore Technology Development

lun wang, M.Eng, B.Eng Offshore Technology Development


INTRODUCTION Keppel’s Research and Development arms, Offshore Technology Development (OTD) and Keppel Offshore and Marine Technology Centre (KOMtech) are developing an improved design of a hydraulic continuous jacking system. The system features continuous operation by using simultaneously running cylinder assemblies. The hydraulic actuators are designed in such a way to provide additional preload capability and overcome the speed limitation of existing jacking systems. With half of the cylinders mounted below main deck, the design of the whole package including jack case structure is compact and requires less deck space. Alternative operation in discontinuous mode is available in case of a cylinder failure and various emergency situations.

BACKGROUND OF HYDRAULIC JACKING SYSTEM DESIGNSGenerally, the hydraulic jacking systems installed on jack-up platforms fall into one of the following two categories:

1) Discontinuous jacking system

2) Continuous jacking system

In discontinuous jacking systems, there is typically one set of jacking cylinders and two jacking yokes – one is connected with the barrel end of the cylinder and secures the hull to the leg, while the other is connected to the rod end of the cylinder in order to move the hull or leg. Each yoke contains an engagement pin, activated by the hydraulic cylinder. The operation is carried out in two stages:

1. Jacking stage: The fixed yoke pins are disengaged; the moving yoke pins are engaged. The jacking cylinders drive the hull or leg up and down.

2. Idling stage: The fixed yoke pins are engaged; the moving yoke pins are disengaged. The jacking cylinders are in return stroke mode.

In discontinuous jacking system designs, the hull or leg will move only during the jacking stage. As a result, jacking operation will be slower and uneconomic.

On the other hand, continuous jacking system designs feature two sets of jacking cylinders: one set in working stroke mode and the other in idle

Figure 1. hydraulic jacking system designs – discontinuous (left), continuous (right)

leg

upper yoke (fixed)

Jacking cyclinder

Jack housing

travelling yoke

hull

engagement pin

leg

Jacking cyclinder (working mode)

engagement pin

hull

Jacking cyclinder (idle mode)

travelling yokes

stroke mode. The yokes are arranged at different elevations, with the idling yokes moving in the opposite direction of the working yokes. As the jacking cylinders change direction for load transfer, intermissions may occur. A system may have four intermissions in a typical jacking cycle and, as such, is considered a complicated control system.

In this continuous jacking system design, the hull or leg moves continuously and will move in both working and idle mode. As a result, jacking operations will be faster and more economical.

OTD HYDRAULIC JACKING SYSTEM DESIGN CONCEPTThe OTD hydraulic jacking system is a continuous type and designed for high-jacking speed and effective usage of hydraulic power. The design concept is based on only two tiers of jacking units as shown in Figure 2, moving in opposite directions – one tier in working stroke and the other in idle stroke mode. A set of valves controlling the hydraulic flow drives the two tiers, working

simultaneously. The jacking speed will be almost twice that of the discontinuous type. The Hydraulic Power Unit (HPU) is utilized more effectively.

GENERAL ARRANGEMENT OF JACKING SYSTEMThe OTD hydraulic jacking system is intended to be installed on self-elevating offshore vessels as shown in Figure 3. A complete system comprises Jacking Units (JU), Hydraulic Power Units (HPU), Local Control Consoles (LCC), and a Jacking Centre Control Console (JCC). The JCC is located in the jacking control room.

Each leg is arranged around with two tiers of jacking units and safety pins as shown in Figures 4 & 5. A jacking unit is mechanically connected to the hull structure by a pair of jacking cylinders, and engage into the leg holes through a jacking claw fitted with an engagement cylinder and pin. Throughout the jacking process, there will always be at least one complete tier of jacking claws clamped onto each leg.

Figure 2. otd hydraulic jacking units


engagement pin

leg

engagement cyclinder

lower jacking cyclinder

Jack case

Jacking claw

upper jacking cyclinder

hull


A special feature of OTD hydraulic jacking system is its high reliability in case of cylinder failure.

Figure 3. overview of a vessel installed with the otd hydraulic jacking system

Figure 4. Jacking units’ isometric view Figure 5. Jacking units front view

A special feature of OTD hydraulic jacking system is its high reliability in case of cylinder failure. The safety pin will be used to secure the hull to the leg in an emergency situation. In this case, the jacking system will operate discontinuously with almost half the speed of normal conditions. The hull will move up evenly. This design provides a complementary safety guard during the windmill installation process.

For hull jacking, the jacking cylinders create hydraulic force to overcome the hull weight; it will be transferred from jacking cylinders to the sea bed through the jacking claws, engagement pins and legs. For leg handling, the load path is similar but in reverse direction. The jacking cylinder mounting is accomplished as pull type with the barrel end pointing downwards and the rod end pointing upwards. In this way, the jack case is excluded from the load path allowing its size to be significantly reduced and the system more compact.

PROTOTYPE TESTINGThe OTD hydraulic jacking system design concept and the operation sequence have been tested with a scaled down, one-leg jacking system prototype in a workshop. The continuous jacking operation (both equal and unequal load conditions) has been achieved by synchronization control.

OTD HYDRAULIC JACKING SYSTEM TECHNICAL SPECIFICATIONThe jacking system’s capacity can be adjusted to satisfy the vessel’s design capacity and specific customer needs. A typical configuration of the system comprises:

Figure 6. otd jacking system prototype

table 1. technical specification of a typical hydraulic Jacking system

technical specification

model otd h2000

Normal Jacking Capacity 2,000 Ton per Leg

Maximum Jacking Capacity 3,200 Ton per Leg

Maximum Holding Capacity 3,800 Ton per Leg

Jacking Speed approx. 900 mm/min

model otd h3000

Normal Jacking Capacity 3,000 Ton per Leg

Maximum Jacking Capacity 4,800 Ton per Leg

Maximum Holding Capacity 5,700 Ton per Leg

Jacking Speed approx. 900 mm/min



CONCLUSIONThe OTD hydraulic jacking system has following advantages:

• High-jackingspeedsystemduetocontinuous operation philosophy

• Effectivesystem–thehydraulicconcept maximises the usage of hydraulic flow and minimises system power requirements

• Simplecontrolsystem–thetwointermissions in typical jacking cycle is simpler than the four intermissions system

• Compactsystem–itrequireslessinstallation space, providing more deck area for wind turbine parts storage and installation operations

• Reliablesystem–ifonecylinderfails,the system can continue to operate at almost half the normal jacking speed and hull will move up evenly.

• Robustsystemforpre-loading,hullholding, leg pulling and in emergency conditions

• Alljackingcylindersmayjackinthesame direction within one stroke

AuTHOR’S CONTACT [email protected]

REFERENCES [1] Emre uraz, “Offshore Wind Turbine Transportation and Installation Analyses”, Master Thesis, Gotland university, Sweden, June 2011

[2] Mikx, Johannes Wilhelmus Jacobus, Dordrecht (NL), “Jacking System for a Leg of a Jack-up Platform”, uS Patent Application Publication No.12/709213, Aug 2010 http://www.freepatentsonline.com/y2010/0215439.html

Simple control system – the two intermissions in typical jacking cycle is simpler than the four intermissions system


Simulating LNG Liquefaction Plant Start-up 25

Komtech has deVeloped its pRopRietaRy lng liquefaction pRocess, pRenex, targeted for the offshore associated gas and small stranded gas fields, with optimisations implemented using HYSYS steady state simulation. Process control and system behaviour is an important consideration for a LNG plant. Dynamic simulation has been adopted to simulate the start-up process with emphasis on the performance of the main cryogenic heat exchanger (MCHE), the most complex piece of equipment in a liquefied natural gas (LNG) plant, to demonstrate that the design is robust. Nitrogen gas is utilised as the refrigerant and its primary source for cooling is by isentropic expansion from a turboexpander. During start-up, the main concern is the possible thermal stress in the material due to large differential temperatures between hot and cold streams in the MCHE. In addition, the rate of which the material is cooled down during start-up should not exceed 60ºC per hour as the flows are being introduced to the MCHE. Dynamic simulation results show with proper setting of the turboexpander inlet pressure to adjust its speed can indirectly minimise thermal stress in the MCHE and furthermore control the cool down rate.

Simulating LNG Liquefaction Plant Start-up

sui Jian Jun, PhD, M.Eng, B.Eng

Jens walleViK, MSc, BSc

sheng xiaoxia, PhD, B.Eng

Kshudi Ram saRKaR, MTech

chong wen sin, CEng, MSc, B.Eng

MULTI STAGE N2 COMPRESSOR

CLEAN NATURAL GAS

VLV-103

E-301

13-3129135

K-100 K-101 K-303 LNG

19

27162

18

MCHE

V-101 FLASH DRUM


INTRODUCTION Liquefaction of natural gas with a volume reduction to 1/600 is of great importance in enabling transport from a supply source to market at lower or even atmospheric pressure. Due to the significant increase in global natural gas demand and environmental concern it is expected that the LNG business will attract more attention from energy industries. Improvements in the gas treatment and liquefaction process have currently enabled the use of LNG where previously it was unviable to monetize fields that were classified as stranded [1].

Many LNG liquefaction plants utilize a mechanical refrigeration cycle such as the mixed refrigerant type where cooling is generally achieved by heat exchange with one or more refrigerants including propane, propylene, ethane, ethylene, nitrogen, methane or mixtures thereof, in a closed loop or open loop configuration. One of the advantages of using a mixed refrigerant cycle is the ability to concoct a multi-component refrigerant that can provide an optimal level of heat transfer efficiency.

To reduce the capital cost of the LNG process with little compromise in specific power consumption, Linde, Kapitza and Claude have proposed expansion devices for gas liquefaction systems. Substantial intellectual property has been developed on this matter. (See references 2, 3, 4, and 5). In the expander based process a stream of gas at high pressure is reduced to low pressure through isentropic expansion, generating refrigeration and

mechanical energy. The low pressure gas is refrigerated and the mechanical energy extracted is used to drive a compressor to partially recompress the gas. The most important equipment in an expander based process is the turboexpander and current developments in the design are resulting in efficiencies greater than 85%. Increasing efforts have been put in to developing new processes in light of the emerging opportunity to challenge the current LNG paradigm.

Steady state HYSYS simulation is widely used for process design and process optimization. However, steady state HYSYS does not have the sufficient capabilities to study transient process behaviours such as during plant start-up and shutdown. Aspen HYSYS Dynamics is a rigorous process simulation engine which is widely used to dynamically model the facility, focusing on the control analysis and system behaviour.

DYNAMIC SIMULATION OF THE PROCESSThe steady state simulation of the LNG plant is constructed based on the process properties provided including temperatures, pressures and mass flow rates of individual streams and equipment. After converting to a dynamic state simulation, equipment data obtained from vendors, such as the type and speed of the compressors, the head and speed of the turboexpander and the different layers and heat transfer efficiencies of the MCHE, is input in order to model dynamic behavior. To facilitate the control of the process, any necessary flow and temperature controllers are installed as

Figure 1. Partial schematic diagram of the lng plant

Figure 2. temperature profile of the two streams (13-3 & 16) with the turboexpander at high rPM

shown in green in Figure 1. After running the dynamic simulation, the transient process is presented in the form of line charts. The mass flow rate of the feed natural gas starts from the minimum of 0 kg/h and eventually reaches the maximum design value. The temperatures at the compressors, the turboexpander and the MCHE are also changing and approaching the designed values. Finally, the dynamic simulation reaches a stable state in which the process properties are closely matched to those in the steady state simulation by ±5%, indicating validity of the dynamic simulation.

RESULTS & DISCUSSIONThe temperature of the turboexpander discharge is the key parameter to be adjusted in order to control the cooling down rate of the MCHE and minimize any possible thermal stress. It is feasible to adjust the suction pressure of the turboexpander with the purpose of varying the speed of the expander which provides a solution to minimize both the

differential temperature and cooling down rate in the MCHE. The temperature difference between stream 13-3 after the turboexpander and stream 16 before the J-T valve is used as the indicator for the temperature difference of the hot and cold streams.

Figure 2 shows the temperature profile of the two streams (13-3 & 16) with the turboexpander/compressor operating at high RPM and Figure 3 displays the temperature profile of the same two streams with the turboexpander/compressor operating at low RPM. It is noticed that the temperature difference in Figure 2 is much larger than in Figure 3, indicating that operating the turboexpander/compressor at low RPM will mitigate the risk of thermal stress when compared to operating at high RPM. For example, close examination of the differential temperature at 5 minutes in Figure 2 shows that it is about 30°C at high speed, approximately twice the difference in Figure 3 at low speed.


-50 -

-55 -

-60 -

-65 -

-70 -

-75 -

-80 -

-85 -

-90 -

-95 -

-100 -

-105 -

-110 -

-115 -

-120 -

-125 -

-130 -

-135 -

-140 -

-145 -

-150 -0 5 10 15 20 25 30 35 40 45

time (min)

tem

per

atur

e (°

c)

stream 16 before J-T valvestream 13-3 after turboexpander


Figure 4. MChe cooling with a nitrogen loop

-50 -

-55 -

-60 -

-65 -

-70 -

-75 -

-80 -

-85 -

-90 -

-95 -

-100 -

-105 -

-110 -0 5 10 15 20 25 30 35 40 45

time (min)

tem

per

atur

e (°

c)

stream 16 before J-T valvestream 13-3 after turboexpander

Figure 3. temperature profile of the two streams (13-3 & 16) with the turboexpander at low rPM

0 10 20 30 40 50 60 70 80 90

time (min)

-50 --55 --60 --65 --70 --75 --80 --85 --90 --95 -

-100 --105 --110 --115 --120 --125 --130 --135 --140 --145 --150 -

tem

per

atur

e (°

c)

Furthermore, the temperature of stream 13-3 at the turboexpander discharge reaches a stable state of approximately -140°C when operating at high speed, but only reaches to approximately -100°C at low speed. With the event scheduler feature in HYSYS Dynamics, it is possible to set the speed of the turboexpander/compressor at a lower RPM until the temperature of stream 13-3 is stable and then to increase the speed to maximize the cooling duty, as shown in Figure 4.

Based on the temperature profiles in the above mentioned figures, it is indicated that by manipulating the suction pressure of the turboexpander to adjust its speed can potentially minimize the risk of thermal stress in the MCHE and also regulate the cooling down rate of the MCHE.

CONCLUSIONIn this paper, the dynamic simulation of an LNG liquefaction plant is investigated to get a better understanding of the operational behaviour and control of the process. Emphasis is placed on the most important piece of equipment, the MCHE, to resolve potential issues including possible thermal cracking due to large differential temperatures between hot and cold streams and control of the cooling down rate of the material in the MCHE. The dynamic simulation results provide a solution to these issues and also useful information for the start-up procedure and operation manual.


REFERENCES [1] A. J. Finn, (2009), Are floating LNG facilities viable options, http://www.costain-floating-lng.com/editorimages/HP0709finn.pdf, accessed in November 2011.

[2] U.S. Pat 3,677,019 Gas Liquefaction Process and Apparatus

[3] U.S. Pat 4,638,639 Gas Refrigeration Method and Apparatus

[4] U.S. Pat 5,916,260 Liquefaction Process

[5] u.S. Pat 5,755,114 use of Turboexpander Cycle in Liquefied Natural Gas Process



Ballast Water Treatment System Selection 31

Ballast wateR has Been identified as a potential path of ecological thReat eVident in the past decade woRldwide. In order to control the transfer of sediment and ballast water, including harmful organisms and pathogens, the International Convention for the Control and Management of Ship’s Ballast Water and Sediments was adopted by International Maritime Organisation (IMO) in 2004 and the Convention will enter into force in the near future. In addition, there are several current and upcoming regional and local regulations to control the discharge of ballast water from ships. For ship owners to comply with the regulations, ships will be required to install and use IMO type approved ballast water treatment system, which can process ballast water to meet discharge standard requirement.

There are several types of available technologies and systems. However, installation of a ballast water treatment system with type approval onboard does not ensure compliance with regulations since each system has its own advantages and limitations. The objective of this paper is to explicate the applicable legal framework at global, regional and national levels regarding ballast water, investigate available technologies and systems to facilitate ship owners to comply with regulations, as well as to address the crucial criteria of ballast water treatment system selection and evaluation.

Ballast Water Treatment System Selection

prapisala thepsithaR, PhD, M.Eng, B.Eng

chong wen sin, C.Eng, MSc, B.Eng

nirmal Raman guRunthalingam, B.Eng



INTRODUCTION Shipping plays an important role in freight transportation and transports 80% of the world’s commodities. Oceangoing vessels require ballast water in order to maintain safe and efficient operations. Ballast water is typically taken onboard a vessel at one port after unloading cargo. This is to maintain its stability and structural strength. When the vessel arrives in another port for cargo loading, ballast water is discharged to offset additional weight caused by the cargo. In some cases, ballast water is taken onboard a vessel in order to travel under bridges safely. It is also discharged to lighten the vessel in order to maintain clearance under the keel when encountering any obstacle in the seareef or when entering to relatively shallow areas such as the navigational channels or berthing areas[1].

Ballast water tanks are normally located at the lowest part of the vessel below the machinery rooms. Types and sizes of ocean-going vessels determine ballast water system of a ship, including a total ballast tank size, a pumping rate and numbers of ballast tanks of the ship. The total capacity of ballast tanks of a ship ranges from 3,000 m3 with the pumping rate of 250 m3/hr for small container ships to 95,000 m3 with the pumping rate of 5,800 m3/hr for ultra large crude carriers [2]. The number of ballast tanks may also range from 10 to 16 tanks.

It is estimated that at least three to five billion tones of ballast water are transferred by transoceanic shipping activities annually [3]. Ballast water and sediments carried by ocean-going vessels containing organisms ranging from bacteria to fish have been well recognised as a potential pathway for transference of various species of organisms, including invasive aquatic organisms and pathogens around the world. The introduction of non-indigenous species to new regions has adverse impacts on the environment, including competing for food between native and non-indigenous species, altering food chain due to no natural predator of non-indigenous species in new regions, changing ambient temperature and so on. These impacts can be irreversible. The ultimate consequences include extinction of native species, ecological imbalance, deterioration of port or ship structure and damages to local industry especially at the coastal waters [4].

In the past decades, there were many evidences of detestable impacts of introducing non- indigenous species to new environment. The transference of European Zebra Mussels to North America displaced native aquatic life and altered habitats resulting in changing ecosystems and fouling all hard surfaces, including infrastructure and vessels. In USA alone, the economic loss has been estimated to be more than US$ 1 billion since 1989. Jelly fishes from North America were introduced to North Sea, Azov Sea and Caspian Sea and they reproduced rapidly. As a result, food chains and ecosystems in the mentioned regions were altered. It caused the collapse of fisheries in Black Sea and Azov Sea in the 1990s. Asian Kelp is currently found in southern Australia, New Zealand, USA, Europe and Argentina. It spreads rapidly, displacing native algae and marine life. In addition, harmful algal blooms, commonly known as red tide, in many regions can cause massive killing of marine life, fouling beaches and ultimately a severe impact on tourism and recreation. Some species may contaminate filter feeding shellfish which, if eaten by humans, can cause severe illness and death [5].

From these severe evidences, several parties, including governments and researchers have put tremendous effort to alleviate the impact of transferring organisms contained in ballast water and sediments resulted from marine industry.

BALLAST WATER REGULATIONS WITH THEIR IMPLICATIONS At the international level, the transfer of ballast water and sediments, containing microorganisms and pathogens is controlled under “the International Convention for the Control and Management of Ship’s Ballast Water and Sediments” adopted by International Maritime Organisation (IMO) in 2004. The Convention applies to all ships operating internationally that carry ballast water. Under Article 1 of the Convention, ships means a vessel of any type whatsoever operating in the aquatic environment and includes submersibles, floating craft, floating platforms, floating storage units (FSUs) and floating production, storage and offloading units (FPSOs) [6].

According to the Convention, ships are required either to perform a ballast water exchange (BWE) set in Regulation D-1 or to meet concentration-based ballast water discharge standard set in Regulation D-2 by using approved system of ballast water treatment (BWT). The application date of the requirement is determined according to ship construction date and quantity of ballast water carried on board the ship as shown in Table 1 [6].

BWE set in Regulation D-1 is considered as a simple procedure for ships. Presently, two primary methods of exchanges are being practiced, namely the sequential method and the flow-through method. The objective is to replace the original ballast water to achieve the IMO exchange criteria (i.e. ≥ 95% water exchange). However, the efficiencies for organism removal are likely to be low, particularly considering the sediments left over in the ballast water tanks [4]. Therefore, the Convention only allows ships to achieve the concentration-based ballast water discharge standard (refer to Table 2) eventually by installing BWT system.

table 1. iMo schedule for ballast Water Management

bAllAst WAter YeAr oF shiP ConstruCtion* CAPACitY before 2009 2009 onwards 2009-2011 2012 onwards

< 1,500 m3 • Before end of 2016: BWT only - - - BWE or - BWT • From 2017: - BWT only

1,500 - 5,000 m3 • Before end of 2014: BWT only - - - BWE or - BWT • From 2015: - BWT only

> 5,000 m3 • Before end of 2016: - • Before end of 2016: BWT only - BWE or - BWE or - BWT - BWT • From 2017: • From 2017: - BWT only - BWT only

table 2. iMo standard of ballast Water discharge (d-2)

As of now, the Convention has not entered into force. It will enter into force 12 months after ratification by 30 States representing 35 percent of the world shipping tonnage. By October 2011, the Convention has been ratified by 30 countries representing 26.44% of the world shipping tonnage [7]. The Convention needs the ratification of two more countries to fulfill the shortage of 8.56% [8]. It is expected that the implementation will be within few years. Due to uncertainty of the enforcement date, it is difficult for ship owners and ship builders to plan for the equipment and utilities required.


organism Category regulation

Plankton, > 50 µm in < 10 viable cells/m3 minimum dimension

Plankton, 10-50 µm < 10 viable cells/ml

Toxicogenic Vibrio Cholera < 1 cfu* / 100 ml (O1 and O139)

Escherichia coli < 250 cfu* / 100 ml

Intestinal Enterococci < 100 cfu* / 100 ml

*cfu = colony-forming unit


In addition to IMO regulations, several regions and countries have already introduced and implemented specific regulations regarding the standard of ballast water discharge. These include California, New York, USA, Australia, Panama and Buenos Aires. Some of the areas, e.g. US Ports plan to have even more stringent standard than the IMO standard while the ballast water discharge is prohibited in some of the areas (e.g. Panama). Owners should check the regulations and requirements of local and regional authorities carefully [9].

AVAILABLE TECHNOLOGIES FOR BALLAST WATER TREATMENT AND DEVELOPMENT TRENDS To comply with the International Convention for the Control and Management of Ship’s Ballast Water and Sediments set by IMO, ships are eventually required to process their ballast water to meet the requirement of discharge standard according to Regulation D-2 by using IMO type approved ballast water treatment system. There are several technologies available for ballast water treatment. In general, there are two major steps for ballast water treatment, i.e. solid-liquid separation and disinfection.

Solid-liquid separation involves the separation of particulate matters, including large suspended microorganisms (>40-50 µm) from ballast water using mechanical devices, particularly filtration and hydrocyclone. Solid-liquid separation is normally used in conjunction with disinfection to enhance overall performance of the system. Concerns over the application of solid-liquid separation processes include pressure drop of ballast water line and waste stream containing suspended solids.

Disinfection is a crucial step to ensure that organisms contained in ballast water and sediments are removed, killed, inactive and/or unable to reproduce. The disinfection technologies can be classified into three major groups as follows:

(1) Physical or physicochemical disinfection:

- Ultraviolet Irradiation (UV)

- Advanced Oxidation Process (AOP) using UV in conjunction with titanium dioxide (TiO2)

- Flocculation and coagulation

- Ultrasonic

- Cavitation

(2) Chemical disinfection:

- Ozonation

- Electrochlorination and electrolysis

- Adding chlorine compounds

- Adding biocides

(3) Asphyxiation:

- Deoxygenation

There are many available technologies in the market, including disinfection mechanisms, major components required, treatment time, treatment positions, consumables, by-products generated, operation-ability, power consumption, footprint, and specific concerns [10-14]. Up to now, there are 8 technologies commonly used as a major component of ballast water systems and these include UV irradiation, AOP (UV/TiO2), ozonation, electrochlorination/electrolysis (EC/EL), chlorine compound injection, biocide/chemical dosing, flocculation/coagulation and deoxygenation. Evaluation has been done and shown that each technology has its own advantages and limitations. In other words, there is no single solution that can fit the needs of every vessel.

For example, the interest in UV irradiation results from its applicability with seawater and freshwater, a short treatment time (i.e. in seconds), no toxic by-products generated during the treatment process and no requirement of chemicals for neutralization before ballast water discharge. However, it is not suitable for water with high turbidity. The treatment using UV irradiation has to be performed twice, i.e. at ballasting and deballasting, to ensure the quality of ballast water discharge. Furthermore, the operating cost is relatively high in comparison with other technologies.

As for EC/EL, it has several unique advantages, especially requirement of the treatment at ballasting only and low operating cost. However, it is not suitable for water with low salinity.

It can generate explosive gas (i.e. hydrogen gas (H2)) and some potential carcinogenic substances (i.e. Trihalomethane (THM)). Before ballast water discharge, neutralization using sodium bisulfite (Na2HSO3) is a necessary step to adjust the value of Total Residual Oxidants (TRO) in ballast water.

Therefore, combinations of technologies have been adopted in some cases. For example, EC/EL system has a small footprint and low power consumption. However, it is not suitable for seawater with a low salinity and fresh water. To eliminate the mentioned limitation, it is used in conjunction with ozonation. It should be noted that the overall system will be more complicated in terms of equipment, piping and control system.

SYSTEM SELECTION CRITERIABy December 2011, there are 22 IMO type approved ballast water treatment systems with reference to Guideline for approval of ballast water management systems (G8) and Procedure for approval of ballast water management systems that make use of active substance (G9) [15,16].

As discussed previously, ships will be required to install and use a type approved ballast water treatment system to process their ballast water to meet the D-2 standard before discharge. It should be noted that even after installing an IMO type approved system, the owner and/or operator are still responsible for compliance of the discharge. It should be emphasized that no single system fits the needs of all vessels in different situations. In other words, installing a ballast water treatment system with type approval does not guarantee that system works well on all vessels in all circumstances. Therefore, it is very crucial for the owners and/or operators to select an appropriate system for their ships. This is to ensure that they will not violate the regulations unintentionally.

When selecting a ballast water treatment system, one must consider several aspects. The parameters to be considered consist mainly of: (1) operation-ability of the system determined by vessel specific information (i.e. ship type and trade route); (2) applicability of the system in terms of installation and its effects to ship operation (e.g. footprint,

power consumption, explosive gases, consumables, etc.); and (3) investment and operating cost.

KOMtech has developed a model to simplify the selection of ballast water treatment system according to the mentioned criteria. The model focuses on operation-ability and applicability. The model also provides an estimated operating cost of each technology. Due to the fact that the investment cost will also be determined by the market mechanisms. Therefore, it was not included in the model. The input of the model includes ballast water pump flow rate, turbidity of water, salinity of water and voyage time. The potential systems of ballast water treatment will be shown with their specification, i.e. estimated size, range of power requirement, consumables, toxic or explosive gas generated, corrosion and waste generated.

CASE STUDYThis section illustrates how to apply the criteria of ballast water treatment system selection to evaluate suitable systems for a certain ship. Table 3 shows ship information used in the evaluation.

After evaluating the data in table 3 (i.e. ballast water pump flow rate, turbidity, salinity and voyage time) potential ballast water treatment systems will be shown as a result of the input data. The potential systems to be applied in this case are as follows:

- Filtration + UV Irradiation

- Filtration + ozonation

- Filtration + EC/EL (side stream)

- EC/EL (full stream)

- Filtration + Coagulation

- Filtration + Deoxygenation

In terms of footprint and power consumption, EC/EL-based systems are more superior to other systems. The ballast water treatment is also required at ballasting only. However, chemicals generated from the system may expedite ballast water tank corrosion. Other concerns include explosive gas monitoring and neutralisation of TRO before ballast water



discharge. The ozone-based technology has a high footprint and relatively high power consumption while UV-based technology has a very high power consumption. In terms of operating cost, EC/EL-based systems are still more superior to other systems. Coagulation system has a large footprint, requires relatively high power consumption and produces large amount of waste. Deoxygenation-based system can retard the corrosion in ballast tank but it has a very high operating cost in comparison with other systems.

CONCLUSIONBallast water from shipping activities has a significant impact on marine ecology and global economy. The transfer of sediments and ballast water, including microorganisms and pathogens are currently controlled by several regional and local regulations. IMO regulations are likely to be enforced in the years to come to alleviate the impact of ballast water. To comply with the regulations, ships are required to install a ballast water treatment system to control the quality of ballast water before discharging. There are several available technologies for ballast water treatment in the

market, including UV Irradiation, advanced oxidation process, electrochlorination/electrolysis, ozonation, coagulation and chemical injection. In many cases, ballast water will undergo solid-liquid separation prior to use in conjunction with the mentioned technologies. Although the ballast water treatment system has been installed onboard, the owner and/or operator are still responsible for compliance of the discharge throughout the vessels life. A certain system might not work in all conditions of ship operations. In addition to investment and operating cost, several technical parameters need to be considered when selecting ballast water treatment systems.

In order to facilitate ship owners and operators to ensure the compliance, KOMtech has been focusing on upcoming regulations with their implications, trends of technology development, available systems in the market and evaluation/selection of suitable systems. In addition, KOMtech has developed a model to simplify the selection of ballast water treatment system and the model will be able to assist ship owners, especially Keppel clients to select an appropriate ballast water treatment system for their vessels.

table 3. ship information – Case study

IMO regulations are likely to be enforced in the years to come to alleviate the impact of ballast water. To comply with the regulations, ships are required to install a ballast water treatment system to control the quality of ballast water before discharging.

ship information

Ship characteristics

Ship type LNG Carrier

Ship size 303 m (L) x 50 m (B) x 12 m (D)

Ballast water

Numbers of pumps 3 pumps (2 pumps for ballasting and deballasting and 1 pump stand by)

Pump rate 3,300 m3/hr (each pump)

Trade route

Voyage Middle east to Europe (7 days per trip)

Salinity >30 psu (no fresh water port)

Turbidity <12 NTu (expected to be normal)

Specific port requirement No


REFERENCES [1] American Association of Port Authorities (2008), Ballast Water Management – Legislative Priorities, http://www.aapa-ports.org/files/PDFs/ballast.pdf, accessed in October 2011.

[2] American Bureau of Shipping (2010), Ballast Water Treatment Advisory, http://www.eagle.org/eagleExternalPortalWEB/ShowProperty/BEA%20 Repository/References/ABS%20Advisories/BWTreatmentAdv, accessed in January 2011.

[3] Lloyd’s Register (2009), Ballast Water Management, https://www.cdlive.lr.org/information/Documents/ShipRight/ShipRight%20update%20V%201.9%20 -%207%20May%2009.pdf, accessed in July 2011.

[4] R. Balaji, and O. B. Yaakob, Emerging Ballast Water Treatment Technologies: A Review. Journal of Sustainability Science and Management, 6:1 (2011) pp.126-138.

[5] North of England P&I Association (2009), Ballast Water Management, http://www.nepia.com/cache/files/1722-1261067181/LPBriefing- BallastWaterManagement.pdf, accessed in August 2011.

[6] International Maritime Organisation (2004), International Convention for the Control and Management of Ships’ Ballast Water and Sediments 2004, http://www.sprep.org/legal/documents/Konvention_en.pdf, accessed in February 2011.

[7] ClassNK (2011), Ballast Water Management Convention, http://www.classnk.or.jp/hp/en/info_service/ballastwater/index.html, accessed in November 2011.

[8] BIMCO (2011), Ballast Water Convention – Looking to Panama, https://www.bimco.org/News/2011/11/03_BWC_looking_to_Pananma.aspx, accessed in November 2011.

[9] Det Norske Veritas (2005), Ballast Water Scoping Study – North Western Europe, http://www.seas-at-risk.org/1mages/North%20Sea%20BWM%20 Scoping%20Study%20(1-2).pdf, accessed in January 2011.

[10] A. S. Stasinakis, use of Selected Advanced Oxidation Processes (AOPs) for Wastewater Treatment – A Mini Review, Global NEST Journal, 10:3 (2008) pp. 376-385.

[11] Great Ships Initiative (2011), Final Report of the Land-Based, Freshwater Testing of the Alfa Wall AB PureBallast® Ballast Water Treatment System, http://www.nemw.org/GSI/GSI-LB-F-A-2.pdf, accessed in November 2011.

[12] B. C. Nielsen, (2006), Control of Ballast Water Organisms with a Seawater Electrochlorination and Filtration System, http://www.fish.washington.edu/research/publications/ms_phd/Nielsen _B_MS_Sp06.pdf, accessed in November 2011.

[13] California Environmental Protection Agency (2002), Evaluation of Ballast Water Treatment Technology for Control of Nonindegeneous Aquatic Organisms, http://www.calepa.ca.gov/publications/Reports/Mandated/2002/BallastWater.pdf, accessed in January 2011.

[14] N.E.I. Treatment System LLC (2011), Ballast Water Treatment – Ballast Tank Protection, http://www.sam-gong.co.kr/english/NEI_brochure_full_v2.pdf, accessed in November 2011.

[15] International Maritime Organisation (2005), Guideline for Approval of Ballast Water Management System (G8), MEPC 58/23, 10 October 2008, http://www.bsh.de/de/Meeresdaten/umweltschutz/Ballastwasser/MEPC_125_(53).pdf, accessed in November 2011.

[16] International Maritime Organisation (2008), Procedure for Approval of Ballast Water Management Systems that Make Use of Active Substances (G9) MEPC 57/21, 4 April 2008, http://www.hakuyohin.or.jp/guideline_G9_rev.pdf, accessed in November 2011.



Wet Scrubbing Process for Marine Emission Control 39

Wet Scrubbing Process for Marine Emission Control

liu ming, PhD, M.Sc, B.Sc.

prapisala thepsithaR, PhD, M.Eng, B.Eng

chong wen sin, C.Eng, MSc, B.Eng

nirmal Raman guRunthalingam, B.Eng

sashikant madgula KRishna, M.Sc, B.Eng


moRe stRingent Regulations on emissions fRom ocean-going Vessels aRe Being legislated in Recent yeaRs. As the need to comply approaches, various technologies and processes are developed and implemented to reduce emission to the atmosphere. This paper outlines the wet scrubbing processes with the emphasis on SOx removal and their technical and economical aspects. A process (Optima 2 Plus) developed by KOMtech has proven to be more cost-effective and achieve more consistent performance.


INTRODUCTION Marine EmissionAs a result of fuel combustion for propulsion and power generating, ocean going vessels produce exhaust polluting emissions including sulfur oxide (SOx), nitrogen oxides (NOx) and particulate matters (PM). These pollutants are implicated in several potential adverse impacts on the environment, including global warming, ozone depletion and acid rain.

Marine Emission RegulationsIn general, any ship operating in an ECA (Emission Control Area) will have to either burn low-sulfur fuel or utilise an EGCS (Exhaust Gas Cleaning System). Emissions from marine vessels are currently controlled by several international, regional and local regulations. At the international level, atmospheric emissions from marine vessels are regulated under

Annex VI of International Convention for the Prevention of Pollution from Ship (MARPOL 73/78) set by International Maritime Organisation (IMO). The regional and local regulations include European Union Commission Directive (EU Directive), the United States Environmental Protection Agency (EPA) and California Air Resources Board (CARB).

ECA geographic scope includes the Baltic Sea, North Sea and English Channel, and will expand to include the North American zone, extending approximately 200 nautical miles offshore in 2012. Zones for Puerto Rico and U.S. Virgin Islands have also been accepted. Eventually, the ECA scope may include all coastal areas of the world. The geographical areas outside of ECA zones are subjected to world-wide limits set by IMO. The current and upcoming ECA Emission Regulation is shown in Table 1.

table 1. global and regional emission regulation (Current and upcoming)

2025

2020

2019

2018

2017

2016

2015

2014

2013

2012

2011

2010

2009

2008

2007

2006

2005

Entered

into force

Alternative date**

Join IM

O

* Emission Control Area (ECA): Baltic Sea and North Sea for SOx and North America (including most of the u.S and Canadian Coast) for SOx and NOx. ** Subject to the feasibility review of fuel availability completed by 2018.

regulation Year

MARPOL Annex VI

- SOx:

a) Global S<4.5% S<3.5%

S<

0.5%

b) ECA* S<1.5% S<1.0% S<0.1%

EU

- SOx:

a) General S<1.5% S<1.0% S<0.1%

b) Port S<1.5% S<0.1%

USEPA

- SOx: S<1.0% S<0.1%

CARB (California)

- SOx: S<0.5% (MDO),

S<1.5% (MGO) S<0.1%

NOXIOUS POLLUTANTS FROM MARINE EMISSION AND THEIR REMOVAL BY WET SCRUBBINGOverview of Control Strategies:In principle, there are two types of strategies for SOx emission control, “throw-away process” and “regenerative process”. In the throw away process, SOx from exhaust gas is captured and converted to a more stable and harmless form before being discharged to environment. And in the regenerative process, SOx is captured, concentrated and recovered as a resource for other value added processes. The regenerative process requires additional setup, storage and reception facilities at port. Hence, regenerative process is not applied on shipboard thus far. Majority of the proposed system for marine SOx removal today belong to the class of “discard process”.

SOx Formation and Removal Mechanisms SOx emission from marine vessel is the result of fuel combustion that contains either organic or inorganic bond sulfur. The majority of sulfur is oxidized into SO2 with small amount SO3 coexisting due to thermodynamic limitation. A minute amount of inorganic sulfur is converted into sulfate particulate that contributes to a visible smoke.

As an acidic component from exhaust gas, the most straightforward way to remove SOx is to use acid-base neutralization, which can be carried out conveniently in any well designed reactor/process. There are well developed land based desulfurization processes making use of lime stone operated either in dry or wet state. Gypsum is produced as the neutralization product and is sold as a raw material to make wallboard to recover a fraction of the operating cost. There are also magnesium hydroxide based desulfurization processes which claim to be less prone to scaling and more efficient than the lime based processes. For offshore EGCS, the alkalinity from sea water is a free and readily available source for SOx removal. Therefore, most EGCS proposed onboard are sea water scrubbing process. The natural alkalinity (bicarbonate) of sea water neutralizes SOx and converts it to bisulfite and sulfite form, after which it undergoes further oxidation and is stabilized into the sulfate form that is a natural component of sea water.

SO2 + HCO3- = CO2 + HSO3

-

HSO3- + HCO3

- = H2O + CO2 + SO32-

2SO32- + O2 = 2SO4

2-

In general, one unit of alkalinity from sea water is able to capture equal molar amount of SO2 from exhaust gas, after which it requires additional equal amount of sea water to neutralize the acidic wash water in order to satisfy discharge standard. For example, every 1 MWh brake power generated from a typical low speed marine engine needs 180 kg of HFO (3% wt sulfur), and in order to capture and stabilize the SOx emission, 147 m3 of sea water with average alkalinity 2300 µmol/kg has to be consumed. However the natural alkalinity of sea water may drop drastically, especially in enclosed shallow water body near fresh water estuaries, the sea water scrubbing process may be too energy intensive or impractical. Therefore, the open loop sea water scrubbing systems will be facing practical challenges in Baltic Sea ECA, where the natural sea water alkalinity could drop to only 500 µmol/kg.

Particulate Matter (PM) (Formation, Effects, Removal method)Along with SOx emission, particulate matter (PM) is another pollutant of concern. PM is a result of heavy fuel combustion. HFO (heavy fuel oil) used mostly by ocean going vessel is a residual fuel generally left over from refining processes which tends to have a high sulfur and metals content. Incomplete combustion, high impurity (sulfur and metal) concentration give rise to the amount of PM generated. Due to the fact that HFO quality varies from different refineries, it is difficult to establish precise relationship between the fuel sulfur content and PM emission [1].

In a wet scrubbing process, PM is removed mainly by impact and diffusion. Particles greater than 1 micron cannot follow the streamlines around liquid droplets as the particle’s mass causes it to break away from the streamlines and hit the droplet. For very small particles (< 0.1 micron) moves randomly as a result of gas molecules’ continuous bombardment, the motion is diffusion in nature and the particle is collected by random colliding with liquid droplet in a confined space. PM removal always happen simultaneously with SOx removal in a wet scrubber when hot exhaust gas contacts scrubbing water and cools rapidly, the particles serve as condensation nuclei and become larger for ease of capturing. There are other methods of PM removal without wet scrubbing such as ESP (electro static precipitation) and acoustic agglomeration; however these equipments have not come into commercial stage in the marine industry.



Although not regulated at the moment by IMO, PM emission needs to be considered in the SOx scrubbing process as it affects the discharge water quality which is regulated by IMO (MEPC 184/59) in terms of PAH (poly aromatic hydrocarbon) and turbidity. Any sea water scrubbing process has to take into account the PM trapped in liquid phase carefully before it is implemented.

REVIEW OF CURRENT TECHNOLOGIES OF WET SCRUBBING PROCESS FOR MARINE EMISSION CONTROLOpen Loop SystemThe open loop system relies exclusively on the natural alkalinity of sea water to scrub sulfur dioxide from exhaust gas as 100% of scrubbing water is drawn from sea and subsequently discharged back after it passes through the system.

In particular, sulfur dioxide removal takes place in a scrubber designed to accept maximum sea water flow, catering for the minimum possible alkalinity and maximum engine load. The internals of the scrubber are usually designed for the exhaust gas and scrubbing water to have maximum contacting surface area to enhance the adsorption. Several types of scrubbers such as spray type, bubble bath, bubble cap, packing column (structural and random packing) or their combinations are generally used. As it is necessary to minimize cost, weight and back-pressure induced by scrubber, a compromise between these factors needs to be reached.

The spent sea water (wash water) passing through scrubber not only carries acidic bisulfite but also

Figure 1. open loop sea water scrubbing system

suspended particles such as particulate matter (PM) from exhaust gas and silt from ambient sea water. Upon leaving the scrubber, the wash water has to be processed before being discharged back to the sea. The wash water is usually pumped or drained to a hydrocyclone separator that induces centrifugal separation of heavy suspended particulates from the water. At the end of the process, the PM is collected from the bottom of the hydrocyclone separator and stored in a sludge tank.

The wash water free from suspended particles is now mixed with equal amount of sea water that brings pH back to neutral (>6.5), which is required by IMO for overboard discharge in most areas.

Advantages:- The system uses only free natural sea water as scrubbing agent, there is no concern on storage of additive chemicals on board for SOx scrubbing.

- Fewer components, control and storage are required compared to other systems.

- Easy for crew to operate and familiarize.

Disadvantages:- Operating in brackish or fresh water area will render the system inapplicable or failure due to the insufficient natural alkalinity.

- Pumping cost remains high due to the large amount of sea water required as scrubbing and diluting agent.

- Discharge of acidic effluent may be restricted in some regions (such as at berth) where there are more stringent local regulations.

exhaust gas

sea water in take

Wash water

Monitoring system

sea watersludge storage

dilution

reheating

scrubber sludge removal

Closed Loop SystemThe principle of closed loop scrubbing system is to use re-circulated scrubbing liquid media with chemical additive (NaOH) to boost the alkalinity to achieve consistent SOx removal performance. Fresh water instead of sea water is used in order to avoid potential scale accumulating in pipelines and scrubber internals. Most of the scrubbing agent is contained in the scrubbing loop with an inline process tank where chemical dosing, fresh water replenishment, pH measurement and control take place.

Evaporation loss of water as a result of heat exchange between exhaust gas and scrubbing agent is inevitable; the re-circulation line is heat-exchanged with another sea water cooling line to reduce this effect. As the closed scrubbing loop keeps accumulating salt (sulfate), particulate and un-burnt hydrocarbon, it is necessary to bleed off small amount of waste water constantly from the process tank. At the same time, make-up water from the ship’s fresh water or portable water system is added to maintain the same volume in process tank. The “bleed-off ” flow is centrifuged to take out heavy particulates and discharged as clean water. Alternatively, it can also be stored in a holding tank and offloaded when ship arrives at port reception facility.

Advantages:- Consistent performance regardless of sea water conditions.

- Lower pumping cost compared to open loop systems as enhanced alkalinity of scrubbing agent allows for a lower flow rate.

- Better separation of suspended solid contaminants (sludge & PM) can be achieved.

- All waste effluent can be stored onboard to achieve zero-discharge.

- Zero acidifying effect to ambient water.

Disadvantages:- Supplying and storing of NaOH onboard requires additional space and special handling.

- Continuous fresh water supply and storage required.

- More components are required than an open loop system.

- More training and equipment familiarization are required for the crew.

Figure 2. Closed loop scrubbing system


sludge removal scrubber

holding tank

Monitoring system

sea water in sea water out

exhaust gas

sludge storage

discharge to sea (optional) heat exchanger

Process tank

naoh dosingreheating


Hybrid SystemA hybrid system integrates the functions of the open and closed loop systems. The system resembles a closed loop system but incorporates additional components to enable it to operate as either an open or closed loop system. At open sea the system operates as an open loop system to save chemical consumption; at port it has the flexibility to function as closed loop system to avoid issues stemming from water quality or port discharge regulations.

The hybrid system has all of the same components that are present in the closed loop system, with the primary distinction being the presence of two wash water treatment devices. As the open loop mode of operation requires 100% of the water to undergo centrifugal separation, it is therefore necessary to have a secondary device large enough for the higher flow rate. Switching from the closed loop mode to the open loop mode requires a change in the functions of certain components. The seawater pump used to provide cooling water to the heat exchanger in closed loop mode becomes the supplier of dilution water in the open loop mode. The heat exchanger is bypassed in open loop mode. The pump used to circulate fresh water in closed loop mode becomes the source of seawater for the scrubber in open loop mode. The mode change also requires change-over from the small volume centrifuge to the large volume cyclone separator.

The advantage and disadvantage of the hybrid system can be summarized as follows:

Advantages:

- Combines the advantages of open loop and closed loop systems to achieve consistent performance and specified regulatory compliance.

Disadvantages

- Requires the most components and extensive training for crew to be familiarized.

Electrochemical ProcessIn principle, the alkalinity for SOx removal can be met from electrochemical process using sea water, which at the cathode surface alkaline substances such as sodium hydroxide is produced. In addition to SOx removal, the electrochemical system also claims

simultaneous CO2 and NOx removal. However, the removal mechanism remains unknown, and more full scale tests have to be carried out to confirm the effectiveness of the technology.

It is still premature to conclude the feasibility of electrochemical process due to the removal mechanism, system reliability, energy consumption and cost concerns.

Advantages

- No chemical needed because alkalinity is produced from sea water by electrolysis.

- Consistent performance regardless of variation in sea water alkalinity

Disadvantages

- Energy intensive electrolysis process adds to power consumption of the system.

- Deteriorating performance due to scale deposition on membrane and electrode from sea water hardness.

- By-products (chlorine and hydrogen) require special handling, storage and post-treatment.

OPTIMA 2 PLUS - A WET SCRUBBING PROCESS DEVELOPED BY KOMtech FOR SOX REMOVAL ONBOARDA high performance and cost effective wet scrubbing system for SOx removal from marine emission are based on the combination of consistent input alkalinity of scrubbing media, proper scrubber design and a responsive control system. Design constraints such as variation of sea water alkalinity, scaling issue, scrubber back pressure and system turn down ratio have to be addressed. KOMtech has developed an “Optima 2 Plus” (O2+) system that addresses these issues effectively. The system does not require fresh water supply to maintain the scrubbing loop. The process can be operated as open or closed loop configuration, and the scale formation arising from sea water hardness is prevented.

Series of lab tests have been performed to demonstrate the potentials of Optima 2 Plus system. Under proper scrubbing configuration, the sea water alkalinity in the process can be enhanced to an extent significantly higher than

Figure 4. electrochemical process

A high performance and cost effective wet scrubbing system for SOx removal from marine emission is based on the combination of consistent input alkalinity of scrubbing media, proper scrubber design and a responsive control system.

Figure 3. hybrid scrubbing system


Process tank

sludge

Monitoring system

naoh dosing

Wash water 1 Wash water 2

sea water

Wash water treatment

unit-1

Wash water treatment

unit-2

sea water

reheating

scrubber

exhaust gas

dillution

sea water

discharge Water treatment

scrubber stage 2

scrubber stage 1

Monitoring system

exhaust gas

electro magnetic

unit 2

electro magnetic

unit 1


its original level without side effect such as scaling and clogging. As shown in Table 2 at the different stages of SOx scrubbing, it was found that at the scrubber internal where the SOx absorbing takes place, the practical alkalinity of seawater can be enhanced to 25,673 µmol/kg or 47,872 µmol/kg when partial or full absorption of SO2 takes place.

Bearing in mind that the required scrubbing water flow rate is reversely proportional to its alkalinity, the Optima 2 Plus system will require significantly less pumping flow as compared to a sea water scrubber, where the average original input alkalinity is only 2,300 µmol/kg. A simulation of scrubbing exhaust gas equivalent to burning HFO of 3.5 and 2.8 wt% sulfur was conducted in our lab scale setup. We found that the flow rate of enhanced sea water can be proportionally reduced as compared to

natural sea water given the same amount of exhaust gas flow (Table 3). Throughout our tests, the scrubber wash water remains bright and clear showing no signs of solid depositing onto the scrubber internals, and the pH of scrubber wash water can be adjusted to 6.50 to satisfy current IMO requirement. The test results also provide the possibility to incorporate wash water neutralization stage into one scrubber unit to further bring down the space requirement for future system installation. Currently, there are on-going development and optimizing of the process including demo plant tests and ship trials.

A case study was conducted based on two popular sea routes; Rotterdam to Saint Petersburg and New York to Rotterdam, with 3 scenarios representing situations when ships travel fully or partially in

table 2. enhancement of sea water alkalinity in sox scrubbing process

* Practical limit where precipitation and scaling from hot seawater takes place

steps reactions

location Achievable alkalinity of reactions of sea water (µmol/kg)*

Mixing caustic soda Sea water mixer 4,600 directly with sea water and scrubber inlet

Partial absorption of SO2

Scrubber internal 25,673

Full absorption of SO2

Scrubber internal 47,872

Full neutralization Wash water tank 25,673 and scrubber

Full aeration Wash water tank > 133,097 and scrubber

HCO3- + OH- = H2O + CO3

2-

Mg2+ + 2OH- = Mg(OH)2

Ca2+ + CO32- = CaCO3

SO2+ CO32- = SO3

2- + CO2

SO2 + Mg(OH)2 = MgSO3 + H2O

SO2 + CaCO3 = CaSO3 + CO2

SO2 + 2OH- = SO32- + H2O

SO2 + SO32- + H2O = 2HSO3

-

SO2 + MgSO3 + H2O = Mg(HSO3)2

SO2 + CaSO3 + H2O = Ca(HSO3)2

SO2 + OH- = HSO3-

HSO3- + HCO3

- = SO32- + H2O + CO2

HSO3- + OH- = SO3

2- + H2O

2SO32- + O2 = 2SO4

2-

CO2(aq) = CO2(g)

table 3. lab scale scrubber test with sea water of enhanced alkalinity for so2 removal from simulated exhaust gas

sea water sea water sox sulfur content exhaust gas Wash alkalinity flow rate concentration equivalent to flow rate water (µmol/kg) (ml/min) (ppmv) hFo (wt %) (l/min) ph

2300 200 650 3.5 15.85 4.92~4.23

4600 100 650 3.5 15.85 5.96~5.75

6900 66.7 650 3.5 15.85 5.60~3.87

2300 200 520 2.8 19.82 5.45~5.20

4600 100 520 2.8 19.82 5.83~5.67

6900 66.7 520 2.8 19.82 6.50~5.72


table 4. Case study - the payback period (years) of optima 2 Plus system installed on container ship (main engine: 16.3 MW; auxiliary engine: 3.6 MW)

sea route optima 2 Plus open loop system Closed loop system

Rotterdam to St. Petersburg (2015)a 1.0 1.6 1.9

New York to Rotterdam (2015)b 4.4 6.4 8.4

New York to Rotterdam (2020)c 0.7 1.0 1.3

a) Whole trip in ECA (Sulfur cap at 0.1%) b) Partial trip in ECA (Global sulfur cap at 3.5% and 0.1% in ECA); c) Whole trip in ECA (Global sulfur cap at 0.5% and 0.1% in ECA)

ECA. As compared to fuel switching and other scrubbing technology, it was found that the Optima 2 Plus system has a lower payback period for all 3 scenarios (Table 4).

CONCLUSIONCurrent wet scrubbing processes to remove SOx from marine engine exhaust gas are based on neutralization processes that use either sea water or fresh water with alkaline dosing to capture SOx. The alkalinity of wet scrubbing media plays important role in determining scrubber selection and operating cost. Enhancement

of alkalinity usually requires the support of fresh water system to avoid potential scaling arising from sea water hardness.

A wet scrubbing process - Optima 2 Plus is being developed by KOMtech for SOx removal from marine emission. The process is able to achieve consistent removal efficiency without being affected by variation of alkalinity from different input sea water. Reduced pumping of scrubbing media allows the Optima 2 Plus system to have a lower payback period compared to open loop and close loop scrubbing systems.

A wet scrubbing process - Optima 2 Plus is developed by KOMtech for SOx removal from marine emission. The process is able to achieve consistent removal efficiency without being affected by variation of alkalinity from different input sea water. Reduced pumping of scrubbing media allows the Optima 2 Plus system to have a lower payback period compared to open loop and close loop scrubbing systems.



REFERENCES [1] “A critical review of ocean-going vessel particulate matter emission factors”, Todd Sax, Andrew Alexis, California Air Resources Board, 11/09/2007

Lorem Dolor Lorem dolor ipsum 49


in this woRK, the adVanced computational fluid dynamics (cfd) simulation appRoach, together with High Performance Computing (HPC) techniques, was applied to simulate water oscillations in the moonpool of a drillship. The effect of water oscillations on resistance and propulsion performance of the drillship were investigated in detail. The added resistance due to a moonpool may reduce the drillship’s forward transit speed significantly. This resistance can be reduced effectively by adopting an appropriate moonpool design to suppress water oscillations. The CFD simulation results correlate well with the MARIN model test data. It is suggested that the CFD simulation approach be used as a reliable conceptual design tool, for more efficient moonpool configurations, so as to suppress moonpool water oscillations and reduce the added resistance of a moonpool.

CFD Simulation of Water Oscillations in the Moonpool of a Drillship 49

CFD Simulation of Water Oscillations in the Moonpool of a Drillship

wang shengyin, PhD, M. Eng, B. Eng

matthew quah chin Kau, PhD, CEng, CMarEng, FIMarEST


INTRODUCTION A drillship is a marine vessel fitted with drilling apparatus and designed to carry drilling platforms out to deepwater locations. According to the ABS Guide for Building and Classing Drillships [1], a drillship belongs to a ship type and is a displacement hull offshore drilling unit. It can be regarded as an adaptation of a standard seagoing ship of mono-hull form, with an additional substructure for a distinctive moonpool from which the drilling operations may be carried out. A drilling moonpool is used to pass drilling equipment into the water from the drilling platform located on the deck of the drillship. In the early 1970s, drillships were typically equipped with two moonpools – a relatively small Remotely Operated Vehicle (ROV) moonpool and a rectangular-shaped moonpool through which drilling operations take place. In current drillship designs, only the rectangular-shaped moonpool is retained, and is further lengthened to allow dual operation in a single moonpool [2].

A drillship’s transit speed is important since the transition between different oil fields worldwide may require considerable time. It is therefore desirable to reduce as much resistance acting on a drillship in transit to increase its forward speed and to reduce its fuel consumption [2-4]. The moonpool may cause a drastic increase in water resistance due to the existence of moonpool water oscillation [5]. This phenomenon of moonpool water oscillation is complicated due to the strong coupling between vessel and moonpool motions [5]. Both vertical oscillation (or piston motion) and horizontal oscillation (or sloshing motion) of the water in a moonpool may be involved, and the coupling between the heave and surge motions of the drillship and the moonpool water oscillations may yield a large-resistance increase. The challenging issues of how to accurately predict the added resistance with a moonpool and how to effectively reduce this added resistance, have sparked the interest of many researchers, and will be discussed in the following literature review.

LITERATURE REVIEWThe problem of water oscillations in a moonpool when the ship is in forward motion in calm water is illustrated in Figure 1. The excitation mechanism

of water oscillations in the moonpool under calm water conditions in transit is due to vortex shedding formed by flow separation at the leading bottom edge of the moonpool [5]. Vortex shedding introduces instabilities in the flow and initiates a transfer of energy from the flow to the disturbance, giving rise to flow fluctuations and added resistance. The vortex arising from the leading edge of the moonpool due to the abrupt velocity discontinuity would then come into contact with the trailing edge of the moonpool opening. The induced pressure fluctuation would spread over the surrounding flow field, leading to a dominant frequency with large energy in what is called the phase locking phenomenon [5]. Energy concentration may be found in the combination of phase locking, with the hysteresis effects associated with peaks in energy concentration at certain frequencies. The water oscillations may exert great forces on the moonpool structure and also affect the vessel’s motion. Given the strong coupling effect between the vessel’s motion and the water oscillations in a moonpool, large vessel motions as well as the large added resistance of a moonpool may be generated. It should be noted that water oscillations in a moonpool might also occur in waves under stationary conditions, which is not covered within the scope of our paper.

There are many empirical, experimental and numerical methods used in different literature [2-16] to predict the added resistance due to water oscillations in a moonpool when a ship is in forward motion in calm water. Damping devices are employed to reduce the excitation and decrease the motions, and this has become the main method employed to lower the added resistance [5]. The relevant study of liquid motion in an accelerating container has a long research history, as shown in the comprehensive literature review in [6].

Figure 1. Water oscillations in a moonpool with a forward motion of the ship

Empirically based methods were popular in past ship design practises. In such methods, the resistance curve predictions were developed with a combination of model testings and previous design experiences [2]. Accurate prediction of the added resistance of a moonpool in calm water was usually difficult given the possible large variations in hull and moonpool designs and the limited availability of test data.

Experimental approaches in model tests have been widely accepted for the study of added resistance of a moonpool in modern drillships [2,3,8] because of the past successes of such methods to predict the ship’s hydrodynamic performance [7]. Such model testing approach uses a scale model to experiment and generate the information, which can then be applied to a full-scale ship. However, a certain degree of empiricism is still necessary to enhance the prediction accuracy of the resistance of a full-scale ship. Moreover, the experimental approaches may be costly and time consuming, and arrive at an insufficient data points with poor repeatability. To alleviate these, it is common for the organisation that operates a ship model basin to use the CFD software to numerically simulate the complicated flow around ships, although such CFD calculations are not allowed to replace the model tests entirely.

Numerical prediction of the phenomenon of water oscillations in a moonpool excited by the forward motion of the ship is usually based on the potential flow theory [9-16]. Linear potential flow theory-based solutions are popular because of their simplicity, ease of use and robustness. Further improvement is however required since the hydrodynamic problem of moonpool water oscillations is usually non-linear in characteristic. The 3D fully non-linear potential theory may

represent a significant improvement over the linear theory. The non-linear effects in free surface waves may limit the resonant responses, which would be seriously over-predicted by the linear potential flow theory due to the relatively small linear potential flow damping [14]. However, the potential flow theory is not sufficient to address the complicated phenomenon of water oscillations in a moonpool because it cannot deal with the significant non-linear effects of flow separation and vortex shedding, which would play a significant role in the excitation mechanism [5] of water oscillations.

The use of complete CFD solutions has begun to emerge in recent times [17]. The fully non-linear CFD simulations have the potential to capture accurately the highly non-linear phenomenon of water oscillations in a moonpool. Furthermore, the CFD simulation approach applies the use of HPC techniques [18] to distribute parallel processes across multiple CPUs, and this speeds up the iterative CFD computation. Hence, the CFD simulation approach has the potential to outperform other aforementioned experimental, empirical, or other numerical methods. However, there has been limited use of complete CFD solutions in the literature reviewed in this paper due to various technical barriers in CFD software tools, lack of necessary high performance computing hardware and/or sound CFD experience. In reference to [5], the first results obtained with CFD software ComFLOW were presented to reproduce model tests and moonpool behaviour observed during sea trials. The results showed a relatively good reproduction of flow separation. It appears that the CFD methods are able to model correctly the global physics surrounding the moonpool behaviour [5]. However, some further developments should be made to include the coupling analysis

The fully non-linear CFD simulations have the potential to capture accurately the highly non-linear phenomenon of water oscillations in a moonpool. Furthermore, the CFD simulation approach applies the use of HPC techniques [18] to distribute parallel processes across multiple CPus, and this speeds up the iterative CFD computation.



of the drillship and moonpool motions. In reference to [4], Son et al. used the CFD code FLUENT to investigate the water oscillations phenomenon inside a recess type moonpool. It was shown that a recess type moonpool might offer some improvements in water oscillations as compared to a rectangular moonpool. Nevertheless, the complicated 3D effects on the moonpool behaviour were omitted [4]. In another study, Alsgaard used the open source code OpenFOAM to study the piston mode resonance phenomenon in a moonpool [14]. The work showed some good quantitative agreements in the resonance frequency, though there remain some accuracy problems due to the simplified 2D numerical simulations.

In the literatures [2,5,19], many devices have been reported to suppress the moonpool water oscillations and reduce the added resistance of a moonpool with the forward motion of a ship. Generally, these devices can reduce either the causes of vortex creation or the consequences of large water oscillations [2,5]. Wedges are the classical solution employed to reduce excitation of water oscillations in the moonpool of a drillship in transit [5] without obstructing the moonpool opening. In reference to [2], both wedges and cut-outs were proposed as the resistance mitigation devices. Four moonpool configurations with wedges and cut-outs were also tested in [3]. It was evident that the configuration with a triangular wedge at the leading edge and a cut-out at the trailing edge may significantly decrease the added resistance. Nevertheless, here remains much to cover on designing the most efficient wedge and cut-out [2]

to date.

Other devices to reduce the cause of vortex creation include a grid of flaps, a large single flap mounted on a hinge, a vertical bulkhead and convergent openings [5]. Elsewhere, viscous damping devices such as flanges, horizontal damping plates, floating lids or mats, baffles and damping chambers, have also been developed to reduce water oscillations. Most of these proposed devices may however be difficult to manufacture, and the devices with moving parts may cause operational problems. In this study, only the relatively simple cut-out at the trailing edge of a moonpool was used to mitigate the moonpool water oscillations.

The objectives of this work are to apply state-of-the-art CFD methods and HPC techniques to predict the added resistance of a moonpool accurately and efficiently, in order to investigate the effect of a cut-out at the trailing edge of the moonpool, and to provide reliable insight into the complicated phenomenon of water oscillations in the moonpool of a drillship to facilitate efficient moonpool design.

SIMULATION METHODSThe CFD Simulation ApproachThe added resistance due to a moonpool is usually deduced from the total resistance of the ship with and without the moonpool, as done in the experimental approach in model tests [2]. In this study, the general-purpose CFD software tool STAR-CCM+ [18] was utilised to apply the CFD simulation approach. This CFD software tool incorporates modern software development technology, state-of-the-art computational continuum mechanics algorithms and excellent user-environment design. Furthermore, it maintains a four-month release cycle [18] to ensure that its users are constantly updated with the latest advances.

In the CFD simulation, a 3D computational domain surrounding the ship was pre-defined. To generate a high-quality volume mesh for the computational domain with an appropriate boundary layer resolution, STAR-CCM+’s advanced automatic meshing technology [18] was adopted. Here, the trimmer meshing model, together with a prism layer mesher, was used. A dissipation zone with an extra active damping effect was defined to fulfil a wave reflection damping filter in the CFD solver of STAR-CCM+ [18], similar to, but more robust than, the filter scheme approach to achieving non-reflecting boundary conditions in [20]. The two-phase air and water flow in the computational domain was modelled by the Volume Of Fluid (VOF) method, which was first developed by Hirt and Nichols [21]. The Dynamic Fluid Body Interaction (DBFI) between the two-phase flow and the ship was tackled by STAR-CCM+’s DFBI module [18], and the induced rigid body motions of the ship may be efficiently simulated by its 6-DOF solver. The Reynolds-averaged Navier-Stokes (RANS) equations [18] were also adopted to investigate the complicated flow dynamical behaviour around

the ship. The implicit unsteady solver of STAR-CCM+ [18] was used to solve the RANS equations in a segregated manner. In addition, Menter’s Shear Stress Transport (SST) K-Omega turbulence model [22] was adopted to capture the turbulent wake effects.

HPC TechniquesDue to the intensive computational work required to achieve a converged CFD solution, state-of-the-art HPC techniques were adopted to fulfil the parallel CFD computation to speed up the CFD computation. In this work, HPC techniques were implemented by the operating system Microsoft HPC Server 2008, together with the CFD software tool STAR-CCM+ [18].

The parallel computing operating system Microsoft HPC Server 2008 was used to manage a cluster of 64-bit machines comprising two HP Proliant Z6000 servers, with a total of four quad-core processors (16 cores) and 128 GB RAM, on both enterprise and private networks with a bandwidth of one Gigabits per second (Gbps). Under the Windows HPC Server 2008 environment, the Message Passing Interface (MPI)-based domain decomposition technique was used for efficient parallel computation. The CFD jobs were submitted from the CFD software tool STAR-CCM+ to the Windows job scheduler using MS-MPI either in batch mode or interactively [18]. Under the present configuration, the two HPC techniques of multiple cores and multiple machines were combined to execute jobs consisting of individual tasks to achieve an even more powerful HPC performance. It was important to scale the simulation size to be in line with the number of cluster nodes being used [18]. The optimal scaling was determined from a ratio of the time to compute and the time taken to exchange data between cluster nodes. In this work, the minimum number of mesh cells placed on each cluster node is about 100,000.

RESULTS AND DISCUSSIONSThe present CFD simulation approach and HPC techniques were applied for the added resistance prediction of an ultra-deepwater drillship after their accuracy and efficiency were validated for a US navy combatant ship model. Comparisons with

the MARIN model test data for the drillship and discussion on efficient moonpool design are further provided in our work.

CFD Validation for ResistanceIn this study, the well-accepted surface combatant ship model David Taylor Model Basin (DTMB) 5415 [23,24] was chosen for the present CFD validation analysis. The ship model DTMB 5415 is a towing tank model representing a modern US naval combatant. The model geometry, which includes both a bulbous bow and a transom stern, as shown in Figure 2, has been adopted by International Towing Tank Conference (ITTC) as a recommended benchmark for CFD validation in resistance and propulsion [23].

The principal particulars of this model are shown in Table 1. In the present CFD validation study, only the bare hull model of DTMB 5415, with a scale factor of 1:24.824 and a specified sinkage of -0.01041 m and a trim of -0.108 degree [24], was chosen, as used in the model tests carried out by Olivieri et al. in [25].

table 1. Principal Particulars of the ship Model dtMb 5415

Particular Full scale Model scale

Length (m) 142 5.72

Breadth (m) 17.973 0.724

Draught (m) 6.15 0.248

Wetted Surface 2972.6 4.861 Area (m2)

Speed (m/s) 9.252 2.097

Reynolds Number 1.40E9 1.26E7

Froude Number 0.248 0.28

Figure 2. hull geometry of the ship model dtMb 5415



Figure 3 displays the computational domain for this problem. Due to the symmetry, only a half domain was used. Figure 4 shows the volume mesh generated by the trimmer meshing model of STAR-CCM+ [18]. Since the trimmed mesh is dominantly composed of hexahedral cells, the total number of unknowns would be reduced. Figure 5 is a local sectional view of the trimmed mesh at the symmetry plane (centreplane or x-z plane [26]). Local mesh refinement was implemented to capture the flow separation around the bare hull. Figure 6 shows the convergence history of the resistance components. The total resistance is dominated by the friction resistance since flow separation for the streamlined hull form can be insignificant. Table 2 shows a comparison of

the measured and simulated resistance values. The CFD simulation discrepancy with respect to the measured data is quite small. Hence, the present CFD simulation approach is able to achieve an excellent correlation with the CFD simulation results and the model test data [24,25].

table 2. resistance of the ship Model dtMb 5415

Particular Value

Experimental Result (N) 44.864

CFD Simulation Result (N) 44.106

Error 1.69%

Figure 3. Computational domain for the ship model dtMb 5415

Figure 4. Volume mesh for the CFd simulation

Figure 5. local sectional view of the volume mesh around the dtMb 5415 hull

Figure 6. Convergence history of the resistance components of the dtMb 5415 model

Total Resistance Pressure Resistance Friction Resistance

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100959085807570656055504540353025201510

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Figures 7, 8 and 9 display a further comparison of the simulated and measured wave elevations at three different wave cuts [24,25]. The present CFD simulation results correlate well with the experimental data in [25]. Figure 10 shows the wave pattern simulated by the present CFD simulation approach. It is remarkably close to the theoretical Kelvin wave pattern [27] for an ideal ship with a single bow pressure point, and is characterised by a divergent wave system and a transverse wave system. Figure 11 shows a comparison of the simulated and measured wave patterns. The wave pattern simulated by the present CFD approach correlates quite well with the wave pattern measured by Olivieri et al. [25].

As a whole, the CFD simulation approach is able to achieve accurate simulation results for the DTMB 5415 model [25]. It thus has the potential to replace the actual ship model tests in the future.

Figure 7. Comparison of the predicted and measured elevations along the cut at y/Lpp = 0.0082


Figure 10. Predicted wave pattern generated by the ship model dtMb 5415

Figure 11. Comparison of the predicted and measured wave patterns generated by the model



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CFD Simulations for a DrillshipCFD simulations were further performed for a drillship with different moonpool configurations. The drillship was developed by Keppel FELS Limited. and the corresponding model tests have been carried out by MARIN, as reported in [28]. The principal particulars of the drillship and the moonpool configurations are shown in Tables 3 and 4. The moonpool configurations of the drillship are shown in Figure 12, and that the drillship without a moonpool is also included as “Moonpool 0” to facilitate the present added resistance computation. The configuration Moonpool 1 is a standard rectangular moonpool. Moonpool 1A is shorter with a larger cut-out angle with respect to the vertical direction while Moonpool 1B is longer with a smaller cut-out angle. A comparison study was carried out to find the best performance of these configurations.

table 3. Principal Particulars of the drillship

Particular Value

Length between Perpendiculars (m) 198

Breadth (m) 35

Static Draught (m) 10

Thruster Motor Power (kW) 4500

Number of Thrusters 6

table 4. Principal Particulars of the Moonpool Configurations

Moonpool opening opening Cut-out Configuration length Width Angle (m) (m) (deg)

1 25.2 12.6 0

1A 25.2 12.6 30

1B 29.4 12.6 23.96

Figure 12. Moonpool design configurations of the drillship

(a) Moonpool 0 (with a moonpool)

(b) Moonpool 1

(c) Moonpool 1A

(d) Moonpool 1B

Figure 13 displays a local sectional view of the volume mesh generated by STAR-CCM+ for each moonpool configuration. Local mesh refinement was produced to capture the free surface and flow separation around the bow, stern and moonpool. Table 5 shows a comparison of the parameters of the drillship deduced from the given static draught

(design draught moulded). The length on waterline (LwL), the length overall submerged (Los) and volume displacement ( ) deduced by the CFD software correlate well with the data used in the model tests [28]. Since the maximum difference of 1.56% is quite small, it can be regarded that the CFD geometry representation is accurate enough.

table 5. Comparison of the deduced Parameters of the drillship

Moonpool Configuration Parameter experiment CFd difference

1 LwL (m) 200.05 199.79 -0.13%

Los (m) 204.22 203.04 -0.58%

(m3) 56151.7 57088.0 1.04%

1A LwL (m) 200.05 199.85 -0.1%

Los (m) 204.22 203.04 -0.58%

(m3) 56208.5 57088.0 1.56%

1B LwL (m) 200.05 199.87 -0.09%

Los (m) 204.22 203.06 -0.57%

(m3) 55945.8 56416.0 0.84%

Figure 13. sectional view of the mesh at the symmetry plane

(d) Moonpool 1B(b) Moonpool 1

(c) Moonpool 1A(a) Moonpool 0 (without a moonpool)



Table 6 shows the calm water resistance of a drillship with a forward speed of 10 knots predicted by the present CFD simulation approach. The percentage increases of the added resistance for moonpool configurations 1, 1A and 1B are 57.17%, 14.34% and 25.93%, respectively. The resistance increases can be consistent with the model test results by Veer and Tholen in [2], in which the added resistance of moonpools are around 30% at low or moderate speed, and can be as large as 100% at high speed. As reported by MARIN [28], the added resistance of the present moonpool configurations is quite significant. Due to the lack of any resistance mitigation device, the added resistance of the drillship with the rectangular configuration Moonpool 1 is the largest. Configuration Moonpool 1A, with a shorter moonpool and a larger cut-out angle, has lesser added resistance than configuration 1B, which has a longer moonpool and a smaller cut-out angle. This could be due to the shorter moonpool experiencing less water sloshing [2] and the larger cut-out angle would lessen the amount of water coming up from the trailing edge of the moonpool opening [5]. Table 7 shows a comparison of the forward speed predicted by the model tests [28] and the present CFD simulations.

table 6. Calm Water resistance of the drillship with a Forward speed of 10 Knots

Moonpool total resistance (kn) Added resistance Percentage Configuration due to a Moonpool (kn) increase

0 509 0 0.0%

1 800 291 57.17%

1A 582 73 14.34%

1B 641 132 25.93%

table 7. speed Prediction for the drillship in Calm Water

Moonpool number of Model tests (kn) CFd simulations (kn) error Configuration thrusters used

1 2 10.96 10.08 -8.03%

1A 2 11.91 11.64 -2.27%

1B 2 11.06 11.44 3.44%

1 6 15.07 14.37 -4.64%

1A 6 16.68 15.90 -4.68%

1B 6 15.26 15.76 3.28%

It can be seen that the present CFD predictions correlate quite well with the experimental predictions. The largest discrepancy of 8.03% with respect to the experimental predictions for the configuration Moonpool 1 may be related to the less-than-normal measurement accuracy due to the strong sloshing, as reported in [28].

Figure 14 displays the convergence history of the resistance components for each moonpool configuration. Due to the moonpool water oscillations excited in calm water by forward ship motion [5], there is a pressure fluctuation surrounding the moonpool. The resulting resistance fluctuation may be enlarged or reduced by the ship’s motion because of the coupling effect between the moonpool water oscillations and the ship’s motion [5]. As shown in Figure 14, moonpool configuration 1A would experience the least pressure fluctuation, while configuration 1B would experience the most significant pressure fluctuation.

Figure 15 displays the wave patterns around the hull of a drillship in calm water with a forward speed of 10 knots simulated by the present CFD

Figure 14. Convergence history of the resistance components Figure 15. Comparison of the simulated wave patterns

(d) Moonpool 1B

(b) Moonpool 1

(c) Moonpool 1A


(a) Moonpool 0 (without a moonpool)

(c) Moonpool A1

(d) Moonpool 1B

(b) Moonpool 1

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oscillations in the moonpools are dominated by the sloshing mode (horizontal oscillation) and that the vertical oscillation is insignificant for the relatively longer moonpools. According to Table 4, the length/width ratios of the present moonpool configurations are equal to or larger than two. In the drillship model tests in [2], it was shown that the dominant mode is the sloshing oscillation when the moonpool length/width ratio increases to two. The standard rectangular moonpool configuration Moonpool 1 would generate the largest vortex, the largest upward mass flux at the downstream wall of the moonpool, and the most significant horizontal oscillation inside the moonpool. The resulting added resistance would be the largest, as shown in the simulation results in Table 6. The relatively simple cut-outs at the trailing edge of the moonpool opening would mitigate the water oscillations and the resulting added resistance may thus become smaller. Moonpool configuration 1A is better than 1B in reducing the added resistance since the water horizontal oscillation inside the moonpool would become less energetic and the dominant water-sloshing behaviour would be less significant due to the shorter length of the moonpool opening and the larger cut-out angle.

CONCLUSION The CFD simulation approach, together with the HPC techniques, has been applied to investigate the water oscillation phenomenon in the moonpool of a drillship in forward motion in calm water. In the validation analysis of a ship model, the resistance and wave elevations predicted by the present numerical method correlate well with the well-accepted model test results. In the CFD simulations of a drillship, different moonpool configurations were studied. The numerical results show that the complicated moonpool water oscillation phenomenon can be accurately simulated and the moonpool, with a shorter length and a larger cut-out angle, would reduce the added resistance significantly since the excitation of vortex shedding can be effectively reduced and the dominant sloshing motion can be significantly mitigated. The CFD simulation results correlate quite well with the available MARIN model test data. It is suggested that the CFD simulation approach, with the recent advances in computer hardware and CFD software as well as the well-honed expertise of skilled engineers, should have the potential to replace the actual model tests for future ship design.

simulation approach. The wave patterns become quite complicated due to flow separation and vortex shedding, which may destroy the transverse wave system in the wake [24]. The wave pattern for moonpool configuration 1 shows that the rectangular moonpool would greatly increase the height of the excited wave, leading to the larger resistance, as shown in Table 6.

Figure 16 shows a comparison of water oscillations for different moonpool configurations with a forward ship speed of 10 knots. It can be seen that the water

Figure 16. Comparison of the simulated water oscillations

(a) Moonpool 1

(b) Moonpool 1A

(c) Moonpool 1B

ACKNOWLEDGEMENTSThis work was supported by the Deepwater Technology Group (DTG) of Keppel FELS . We are grateful towards Mr Edwin Nah Hock Choon, (Senior R&D Manager) of DTG, for sharing his insights and expertise in deepwater technology. We would also like to express our sincere gratitude to CFD expert, Mr Peter Ewing, Director of CD-adapco Australia, for his technical support and guidance.


REFERENCES [1] “Guide for Building and Classing Drillships”, American Bureau of Shipping (ABS), Houston, uSA, 2011.

[2] R. van Veer and H. J. Tholen, “Added Resistance of Moonpools in Calm Water”, Proceedings of the ASME 27th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2008-57246, Estorial, Portugal, 2008.

[3] G. S. Virk, H. Chiu, D. R. Deter, C.v. d. Stoep, “Design of the Dynamic Positioning System for the Drillship Glomar C. R. Luigs”, OTC 11954, Offshore Technology Conference, Houston, uSA,2000.

[4] H.-J. Son, S.-H. Choi, M.-W. Kim, and S.-M. Hwangbo, “Drag Reduction of Recess Type Moonpool under Vessel’s Forward Speed”, OMAE2008-57118, Proceedings of the ASME 27th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2008, Estoril, Portugal, 2008.

[5] G. Gaillarde and A. Coteleer, “Water Motion in Moonpools - Empirical and Numerical Approach”, ATMA - Association Technique Maritime et Aeronautique, Paris, France, 2004.

[6] H. N. Abramson, “Dynamic Behavior of Liquids in Moving Containers with Application to Space Vehicle Technology”, NASA-SP-106, 1966.

[7] V. Bertram, “Practical Ship Hydrodynamics”, Butterworth-Heinemann, Oxford, 2000.

[8] J. T. Madhani, “Sloshing Motion of Water in a Moonpool”, Masters by Research thesis, University of Strathclyde, 1985.

[9] C. P. Pesce, E. A. Tannur, and L. Casetta, “The Lagrange Equations for Systems with Mass Varying Explicitly with Position: Some Applications to Offshore Engineering”, Journal of Brasilian Society of Mechanical Science and Engineering, 28:4 (2006) pp. 496-504.

[10] O. M. Faltinsen, “Sea Loads on Ships and Offshore Structures”, Cambridge University Press, Cambridge, UK, 1990.

[11] K. Fukuda, “Behaviour of Water in Vertical Well with Bottom Opening of Ship, and Its Effect on Ship Motions”, Journal of the Society of Naval Architects of Japan, 151 (1977) pp. 107-122.

[12] B. Molin, “On the Piston and Sloshing Modes in Moonpools”, Journal of Fluid Mechanics, Cambridge University Press, 430 (2001) pp. 27-50.

[13] Y. Wei, J. Yang, G. Chen, and Z. Hu, “The Research of Moonpool Size Effect on the Hydrodynamic Performance of FDPSO”, Proceedings of the ASME 2011 30th International Conference on Ocean, Offshore and Arctic Engineering, OMAE1011-49586, Rotterdam, Netherlands, 2011.

[14] J. A. Alsgaard, “Numerical Investigations of Piston Mode Resonance in a Moonpool Using OpenFOAM”, M.Sc. thesis, Norwegian University of Science and Technology, 2010.

[15] C. Maisondieu and P. Ferrant, “Evaluation of the 3D Flow Dynamics in a Moonpool”, Proceedings of the Thirteenth International Offshore and Polar Engineering Conference, Hawaii, USA, pp. 493-500, 2003.

[16] P. Ferrant, “Fully Non-Linear Interactions of Long-Crested Wave Packets with a Three-Dimensional Body”, Proceedings of 22nd ONR Symposium on Naval Hydrodynamics, Washington, USA, pp 403-415, 1998.

[17] ISSC, “Report of Technical Committee I.2: Loads”, Proceedings of the 17th International Ship and Offshore Structures Congress, Seoul, Korea, 1 (2009) pp. 127-210.

[18] “User Guide: STAR-CCM+ Version 6.02.009”, CD-adapco, 2011.

[19] R. A. Ibrahim, “Liquid Sloshing Dynamics: Theory and Applications”, Cambridge University Press, 2005.

[20] K. M. T. Kleefsman, “Water Impact Loading on Offshore Structures - A Numerical Study”, Dissertation, university of Groningen, Netherlands, 2005.

[21] C. W. Hirt, and B. D. Nichols, “Volume of Fluid (VOF) Method for the Dynamics of Free Boundaries”, Journal of Computational Physics, 39 (1981) pp. 201–225.

[22] F. R. Menter, “Two Equation Eddy Viscosity Turbulence Models for Engineering Applications,” AIAA Journal, 32:8 (1994) pp. 1598-1605.

[23] “Benchmark Database for CFD Validation for Resistance and Propulsion”, ITTC QM 7.5-03-02-02, 22nd International Towing Tank Conference, 1999.

[24] L. Larsson, F. Stern, and V. Bertram, “Benchmarking of Computational Fluid Mechanics for Ship Flows: The Gothenburg 2000 Workshop”, Journal of Ship Research, 47:1 (2003) pp. 63-81.

[25] A. Olivieri, F. Pistani, A. Avanzini, F. Stern, and R. Penna, “Towing Tank Experiments of Resistance, Sinkage and Trim, Boundary Layer, Wake, and Free Surface Flow Around a Naval Combatant INSEAN 2340 Model,” Iowa Institute of Hydraulic Research, The University of Iowa, IIHR Report No. 421, 2001.

[26] “Dictionary of Ship Hydrodynamics”, ITTC, 2011.

[27] W. Bascom, “Waves and Beaches: The Dynamics of the Ocean Surface,” Anchor Press, New York, 1983.

[28] J. H. Allema, and G. Radstaat, “Drillship DS 12000: Calm Water Powering Tests”, Report No. 23816-1-DT, MARIN, The Netherlands, 2010.

[29] J. Liang, J. Y. Cao, and C. H. Tan, “CFD for Prediction of Current Load on Drillship”, KOMtech Technology Review (2010). Pg 69 – 74.



New Painting Technologies for Productiivity Improvement 63


pRoductiVity is a measuRe of the efficiency of pRoduction woRK, and can be contributed by man-hours, costs, and quality of work, amongst many other factors. Legislative and commercial drivers constantly challenge productivity in ship and rig-building, as well as in the repair industry.

New regulations in the hiring of foreign workers in the local marine industry, equates to higher manpower costs in the yards. With the new regulations coupled with the limited land area of the yards, the offshore and marine industry needs to develop innovative ideas to increase productivity, and to complete all newbuilds in the quickest turn-around time, at the same time without compromising safety.

Keppel Offshore and Marine (Keppel O&M) takes a serious stance towards productivity in all its yards. This paper discusses how technology is employed in the production activities of the yards to augment productivity of the painting process. The painting projects presented are designed for the painting of the side-shell surfaces of a jack-up rig, at the block erection stage.

New Painting Technologies for Productivity Improvement

yap wee leong, B. Eng

charlie chan chun ta, B. Eng

queK choon Kiat, M.Sc, B. Eng

ng chee yen, B. Eng

wu wenjin, B. Eng


INTRODUCTION The need to improve on the productivity and efficiency in Keppel O&M yardsKeppel FELS has a busy and fast-tracked schedule to build high-specification rigs to meet the customers’ orders. To achieve this with the limited space available in Singapore and limited manpower resources, Keppel FELS has to innovate methods to increase the productivity and efficiency of the production activities. One of the identified processes that can be improved is the painting process during the on-block stage.

Background of painting process in Keppel FELSBefore the hull of a jackup is erected, they are assembled as “blocks”, with stiffeners for strengthening purposes. Blocks forming the external of the hull, known as “side shells”, have vertical flat surfaces, while blocks forming the bottom half of the hull will have flat bottom surfaces. It is thus possible to have blocks with a flat bottom surface and one or two vertical flat walls; or to have a block with no bottom flat surface, and no external flat vertical walls if this block forms the central top block of the hull.

Before assembling the blocks to form the hull, one of the stages require that the blocks undergo grit blasting and spray painting in the paint shop.

Each block is moved to the blasting/paint hall for the blasting and painting processes. Staging is set up to prepare for the blasting process. Spray painting is done on the same staging, after the completion of the blasting process. Both processes are done manually, without any form of jig-type assistance or automation. Workers are required to paint each surface of a block.

The current manual spray method is done using spray guns. The paint is pre-mixed in a separate mixer before feeding to the nozzle via a hose for usage.

The workers will point the nozzle to the unpainted area and spray paint the surface by repeatedly swaying the nozzle in patterned movements. This results in paint wastage, due to:

• Lowtransferefficiency.Thisisattributedto over spray from the worker.

• LargevariationinDFT(DryFilmThickness) leading to redundant coat (i.e., locally higher DFT than the required limit). This is attributed to inconsistent manual spray.

• InsufficientDFT(i.e.,locallylowerDFTthan the required amount). This is also attributed to inconsistent manual spray.

Due to the above differences, re-work jobs are required for a better quality of paint surface which leads to additional manhours and resources for a job.

The objective in this paper is to introduce methods to reduce paint wastage and improve paint coating quality, as well as to reduce manpower for the same job.

DESCRIPTION OF THE METHODCollaborating with Keppel FELS and local vendors, KOMtech engineers has designed specific systems for the painting of the under-hull surface and the side-shell surface of the blocks. The systems are mobile and able to accommodate different hull surface heights and spray paint width.

The advantages of the systems are:

• Overallsetupandpainttimeisshortened

• Consistentdistancebetweenthenozzleand the surface can be maintained, thus reducing the chances of under- or over-spraying, which results in poor DFT quality.

• Amoreworkerfriendlysolution

With the space constraint in Singapore and limited manpower resources, Keppel FELS has to innovate methods to increase the productivity and efficiency of the production activities.

RESULTS AND EXPECTED RETURNSInitial trials on the painting systems have been carried out. Simulated environments were built for the initial trials. To be realistic, actual production workers were employed in the trials. The initial trials show that the paint spray width is wider than that of conventional spraying with a consistent Wet Film Thickness (WFT) in the acceptable range of 250 – 350 microns. Dry Film Thickness (DFT) of the paint was taken after 24 hours, and it was found to be in the acceptable range of 150 microns. Calculations and trial measurements show that the new equipments have the potential to achieve paint savings of about 20%.The new system also requires one less man to complete

ACKNOWLEDGEMENTS

The project is a collaboration between Keppel Offshore and Marine Technology Centre (KOMtech) and the Process Excellence and Painting Sections of Keppel FELS. The authors would like to thank the engineers’ effort in working tirelessly to make the project a success.


the same painting job. In addition, the amount of re-work for the under and over-spray areas were reduced leading to further savings in man-hours and paint.

CONCLUSIONBoth painting projects reduced the number of man hours required for the painting of the hull surfaces of the blocks. More workers can be put to other more urgent work, or the employment and reliance of contract workers can be reduced, thus saving manpower costs. With automation comes consistency, this will also reduce the chances of re-work. On top of this, there can be a significant amount of paint saved as paint wastage is reduced.

New Painting Technologies for Productiivity Improvement 65


Fatigue Assessment of 3-Chord Jacket Substructure for Offshore Wind Turbines in the North Sea 67


JacKet-type suBstRuctuRes foR offshoRe wind tuRBines haVe Been found to Be subjected to a highly dynamic interaction between the structural assembly comprising rotor, nacelle, tower, substructure and soil, and the environmental conditions of current, waves and turbulent wind. As a result, jacket-type substructure designs are fatigue-driven. The Damage Equivalent Load (DEL) simplified method of fatigue life calculation, widely used in the industry, contains simplified assumptions that may lead to overdesign. This paper describes a study carried out on the fatigue assessment of a 3-chord jacket substructure designed for 40m water depth in the North Sea. Taking the load time histories from time-domain simulations carried out with the specialised software BLADED as a starting point, the hot spot stress time histories are calculated with a self-developed software tool and used to compute the fatigue life of structural details, finding that the DEL method can be conservative and non-conservative.

Fatigue Assessment of 3-Chord Jacket Substructure for Offshore Wind Turbines in the North Sea

mariano otheguy, PhD, MSc

yilmaz salman, MSc

wouter henstRa, MSc


INTRODUCTION Since the first offshore wind farm was installed in 1991 off the Danish coast, the offshore wind industry has seen a very rapid expansion [1]. This expansion has been remarkably steep in the last 10 years and is expected to continue over the next decade (Figure 1), being characterised by ever larger turbines installed in harsher environments and also to larger wind farms, water depths and distances to coast (Figure 2). As seen in Figure 2, a large number of wind farms are under construction and consented in the range of 20-50m water depth. Under these conditions, space frame (“jacket”) structures have proven to be fit for purpose and increasingly cost effective, especially for water depths of 30m and beyond [2].

Current innovation trends in offshore wind equipment and structures aim at increasing energy yield and reducing costs. Being the present cost of supporting substructures approximately one fourth of the lifecycle costs of a wind farm and approximately equal to its lifecycle operating cost [3]

(Figure 3), improvements in the construction and maintenance of jacket supporting substructures represent a significant opportunity to reduce the final cost of offshore wind energy.

TIME-DOMAIN SIMULATIONSJacket structures have been successfully used in the oil and gas sector for decades, creating an expertise which is now useful for the development of offshore wind supporting jackets. The key difference between jacket structures used in oil and gas platforms and those supporting offshore wind turbines (OWTs) is the nature and the level of the loads they need to withstand. While wave loads determine the design of oil and gas jacket type fixed platforms, most of the total loading on jackets for offshore wind turbines is wind driven, with load levels exceeding those borne by traditional oil and gas fixed platforms [4].

The oil and gas industry design standard for dynamic loading and fatigue behaviour is based on frequency domain calculations. Unfortunately,

Figure 1. Projected cumulative offshore wind capacity (eWeA and national renewable energy action plans (nreAPs) (source: eWeA)

In a time domain analysis, a large number of computer simulations of short-time operation periods (typically 10 minutes) are carried out on a simplified finite element (FE) model of the whole structural set turbine-tower-substructure. Following standardised design load cases [5] [6], load time histories are produced at certain nodes of the structure, which are post-processed, weight-averaged according to the load case probability and extended to the total service life in order to obtain the final fatigue life of critical structural details.

FATIGUE LIFE CALCULATION METHODSThe fatigue life calculation can be done according to the following methods:

• Damageequivalentload(DEL)

• Sequentialanalysis(SA)

• Fullyintegratedanalysis(hotspotstresstime histories, HSSTH)

Figure 2. distance and depth of online, consented and under construction offshore wind farms (source: eWeA)

this approach is no longer valid for the nonlinear dynamic loading transferred from the rotor through the turbine tower. In order to keep the blades at their aerodynamic optimum performance as well as to avoid resonance in the system, a programmable algorithm controls the individual pitch of each of the blades in real time, introducing complex non-linearities that are best examined with time-domain analyses.

Figure 3. indicative cost breakdown corresponding to the lifecycle costs of an offshore wind farm (3 MW turbines, park size constrained by grid capacity [3])

Foundation, 24%

electrical infrastructure, 24%

Wtg & installation 19%

others, 3%oPeX, 31%



The DEL method, typically used for preliminary design, provides a theoretical load which exerts a fatigue damage which is equivalent to that of a full given time-domain life cycle. This load is calculated from the load time histories (not stresses) assuming a single slope S-N curve (usually slope = 4 is used as a compromise) and is used to calculate a corresponding hot spot stress and S-N curve-based fatigue life calculation of structural details (see flow diagram in Figure 4). However, usual log-log S-N curves are bi-linear, featuring two regions with different slopes (usually 3 and 5). Hence, the DEL method is associated to an uncertainty about which S-N curve slope to be used. Additionally, the DELs must be calculated separately for each load component, neglecting the effects of the actual load combination on the final instantaneous load amplitude, which may lead to self-cancelling but also strengthening effects.

The SA method (which can also be used in combination with DELs) was developed as a natural step linking well known and reliable software packages used for offshore engineering on one hand and for wind turbine load analysis on the other. This method consists of a combination of separate calculations regarding wave and wind loads, which addresses the fatigue analysis with a semi-integrated model. First, an initial offshore engineering FE analysis of the substructure is carried out to compute the structure response to wave loads. A set of matrices and a wave loading history vector are then transferred into the wind load module, where the wind loads are simulated and the corresponding load time histories at the tower-substructure interface are calculated. These load time histories are extended from the tower-substructure interface into the substructure nodes, where the local wave loading is applied. The resulting load time histories now include wind and wave loads and are used for fatigue life calculation based on hot spot stresses and S-N curves. This method splits the system into substructure and rotor-nacelle-tower, therefore it lacks dynamic coupling between these two structural systems

which are connected in reality. For this reason, the sequential method is also regarded as conservative [7] (i.e. likely to provide fatigue damage significantly exceeding that occurring in reality).

Recent software solutions address the time domain simulations of wind and wave loading on a fully integrated model, capturing the structurally coupled nature of an installed and up-and-running OWT. One of these solutions is BLADED, developed by GL Garrad Hassan, which has been chosen for the fatigue analysis of a 3-chorded jacket design by Keppel. BLADED takes a preliminary version of the jacket designed with the DEL method as an input for a set of time-domain simulations which integrates wind and waves. These are applied to a FE model of the entire structural set, which comprises a linearised soil stiffness matrix, beam FE model, wave and wind sub-models and a programmable blade pitch control algorithm [8]. The load time histories obtained are used to calculate their corresponding hot spot stress time histories (HSSTH), using an in-house software tool developed by Keppel (in yellow in Figure 4). These HSSTHs are rainflow-counted and the fatigue damage is eventually calculated as the weighted summation of the damage exerted by each simulation and extended to the entire service life of the system.

The three fatigue assessment methods mentioned above work together with a hot spot/S-N curve approach, standard in the industry [10]. This approach uses parametric or FE calculated stress concentration factors to compute the stress at key hot spots around tubular welds, assessing their fatigue life by means of standardised S-N curves which relate cyclic stress ranges with the associated fatigue capacity according to the Palmgren-Miner rule of linear accumulated fatigue damage.

Certain overdesign is acknowledged in the industry, especially using the DEL method and also the sequential method. This research work is aimed at pursuing a reliable design and also a reduction in production and maintenance costs by means of using and refining the HSSTH method.

Certain overdesign is acknowledged in the industry, especially using the DEL method and also the sequential method.

Figure 4. Flow diagrams of fatigue analysis methods, from left to right: damage equivalent load / sequential method [9] / Fully integrated method (hssth; in-house developed tool in yellow)

del method sequential method

offshore engineering

software

Wind turbine simulation software

Fully integrated method

time-domain simulations

calculation of initial deflection

shapes + stiffness, damping and mass matrices and wave

loading history

wind-only time-domain simulations

load time-histories applied to tower

inteface and extended to the

substructure nodes

calculation of wave loads if necessary

Rainflow counting of

load time-histories

calculation of wave dels if necessary

calculation of dels

local wave loads applied to

substructure nodes

calculation of hot spot stresses

at joints


hot spot stresses at joints

s-n curve s-n curve

fatigue life fatigue life

fatigue life

s-n curve


hot spot stresses time-histories

calculation of hotspot stress time-histories

at joints

wind and waves time-domain simulations

b)



3-CHORDED JACKET: RESULTSFollowing previous in-house work in OWT jacket design [2], an initial design of a 3-chord jacket was produced by Keppel for an offshore location at 40m water depth. The design was subjected to a first preliminary assessment following the DEL method.

The design parameters were completed with generic metocean data corresponding to the German sector of the North Sea, data from the standard NREL 5MW wind turbine [11], DNV design standards [5] and recommended practices [10]. A set of DELs from the turbine dataset was used to compute the fatigue damage corresponding to wind loads, and waves corresponding to the available metocean data were applied to the structure using the software package SAP2000.

After a first design iteration the first natural frequency of the system was kept between the blade passing frequency and the rotor frequency, the two main dynamic excitations in the system. This is important so as to ensure that structural resonance is kept to a minimum. Scantlings were set to comply with the allowed fatigue life of 25 years, leading to a final weight comparable to that of competitive designs currently found in the market. Examining the fatigue life and the ultimate load state (ULS) utilisation (estimated maximum ultimate load divided by the maximum allowed load) of the jacket joints, it was confirmed that in all cases the fatigue response drives the design, being the ULS utilisation always below 25%, hence used as a secondary check. The resulting preliminary design was subjected to time domain simulations using BLADED (Figure 5), the output of which is, at the time this article is being written, subject of on-going work towards its full post processing following the HSSTH fatigue assessment method.

The transition piece, connecting the turbine tower with the jacket, is a key structural assembly bearing very high overturning moment loads. A basic coarse model of a provisional design of the transition piece was used in the BLADED software to compute the time domain simulations (see Figure 6). Its in-depth design and structural analysis under FLS and ULS was carried out with a dedicated detailed FE model in ANSYS.

Figure 5. blAded model of the structural set turbine-tower-jacket in 40 m water depth

Figure 6. blAded render of the Fe model of the jacket piled to the seabed (the transition piece shown is a provisional design for time-effective calculations)

ANALYSIS RESULTSA fatigue life analysis was carried out on the chord-brace connection shown in Figure 7. The DEL and HSSTH methods were applied and an additional HSSTH analysis on the separate load components was carried out, obtaining the final fatigue damage as that exerted by the axial and bending loads separately and summed up afterwards. This additional calculation (which replicates the


Figure 7. structural detail of the jacket

DEL principle of calculating the fatigue damage corresponding to separate load components) enables in-depth analysis of the distribution of stress ranges and number of cycles.

The obtained results show that the HSSTH method, applied on the instantaneous vectorial combination, does not always provide lower fatigue damage than the DEL method. While some hot spots show lower values of fatigue damage, other hot spots in the same joint show similar or even higher values (see Figure 8).

The HSSTH method, applied on the instantaneous vectorial combination, does not always provide lower fatigue damage than the DEL method. While some hot spots show lower values of fatigue damage, other hot spots in the same joint demonstrate similar or even higher values.

Figure 8. normalised fatigue damage at hot spots 1, 5 and 10 of the joint shown in Figure 7, corresponding to a service life of 25 years, calculated with the hssth and del methods

100%

90%

80%

70%

60%

50%

40%

30%

20%

10%

0%hot spot 1 hot spot 5 hot spot 10

hssth method hssth method - separate load components

del method - slope = 4 del method - slope = 5


As shown in Figures 9, 10 and 11, the vectorial combination of axial and bending moments at every time-step leads to a much lower number of cycles (up to 50 times lower) at low stress ranges (below 10 MPa), indicating a lower fatigue damage at low stress ranges. However, due to the logarithmic characteristic of fatigue damage, differences in the distribution of stress ranges and number of cycles in the mid range of both variables affect fatigue damage much more severely than at low stress ranges. For example, Hot Spot 1 exhibits virtually the same HSSTH fatigue damage for separate and for combined loads because the large difference at low stress range level mentioned above (bottom right of graph in Figure 9) is entirely compensated in terms of fatigue damage by an apparently much lower difference between 10 and 30 MPa, where 105-5*106 cycles are counted (highlighted region in the same graph).

Figure 9. distribution of stress range bin and number of cycles corresponding to hssth calculation of combined and separate load components for hot spot 1 of the tubular joint from Figure 7: the resulting fatigue damage in both cases is very similar as seen in Figure 8; the highlighted region shows the area where the vectorial combination leads to clearly higher stress ranges, compensating for the lower stress ranges recorded below 10 MPa (bottom right side of the graph) – data in Figure 9, Figure 10 and Figure 11 corresponds to a service life of 25 years

At Hot Spot 5 this difference is much larger (see highlighted region of graph in Figure 10), resulting in a notable increase in fatigue damage for instantaneous combined loads compared to that of separate loads. As described above, the HSSTH analysis on separate loads gives approximately the same result as the DEL method for the “most appropriate slope value”, indicating that in this case this value is 5.

However, the DEL method ignores the effects of the instantaneous vectorial combination of loads on the final value of the fatigue damage. These effects are highly nonlinear due to the logarithmic nature of fatigue damage, the way the cycle count process works and the fact that the mean stress levels are disregarded. The mean stress level, which is independent of the stress range, is ignored as a common engineering practice, due to the uncertainty of residual stress

number of cycles

hssth separate components

hssth

1 10 100 1000 10000 100000 1e+006 1e+007 1e+008 1e+009

100

10

1

0.1

0.01

str

ess

ran

ge

bin

[M

Pa]

Figure 10. distribution of stress range bin and number of cycles corresponding to hssth calculation of combined and separate load components for hot spot 5 of the tubular joint from Figure 7: the highlighted region shows the area where the vectorial combination leads to clearly higher stress ranges, finally resulting in a higher fatigue damage than that calculated from separate load components


level at weld heat affected zones [10]. Hence, the calculated fatigue damage will depend only on stress ranges, i.e. high mean stress level with low range cycles will exert less fatigue damage than low mean stress level with high range cycles. As a result, the summation of cyclic signals which are in principle independent from each other effectively leads to a new signal with an unpredictable cycle count result, compared to that obtained from the same signals separately and summed up afterwards.

The effects described above explain the notably high fatigue damage found at hot spot 5, beyond that of DEL slope = 4 (highlighted region of graph in Figure 10), showing that the practice of using slope = 4 in order to account for the uncertainty of slope value, regarded as safely conservative, can be actually non-conservative (i.e. underestimating high fatigue damage values).

At some hot spots, the HSSTH calculation on combined loads gives lower values of fatigue damage than that carried out on separate loads. Figure 11 shows that at Hot Spot 10 of the analysed joint, the distribution of stress ranges and cycle number corresponding to separate loads exceeds that calculated from the same loads combined along nearly all the stress spectrum.

ON-GOING WORKThe authors are currently working towards the post-processing of the complete set of load time histories supplied by BLADED. This post-processing includes the calculation of hot spot stress time histories for all nodes in the substructure and all design load cases, and the final calculation of the fatigue life of all critical hot spots in all joints of the substructure. An in-house software tool has been developed and is now under validation. It will also be used to assess the fatigue life of the transition piece.

number of cycles


hssth

str

ess

ran

ge

bin

[M

Pa]

1 10 100 1000 10000 100000 1e+006 1e+007 1e+008 1e+009

100

10

1

0.1

0.01


Figure 11. distribution of stress range bin and number of cycles corresponding to hssth calculation of combined and separate load components for hot spot 10 of the tubular joint from Figure 7: the separate components exhibit higher stress ranges for most of the sample, leading to higher fatigue damage as seen in Figure 8.

This detailed integrated analysis will enable the most accurate fatigue damage calculation currently possible, contributing to avoid overdesign and also to evaluate high fatigue damage where other methods are being non-conservative. It will also enable useful post-processing features as direct comparison between wind plus waves and wind-only loads and identifying the distribution of design drivers in the substructure, enabling selective structural optimisation.

CONCLUSIONIn the context of the current offshore wind farm development, innovation in supporting substructures has a potentially significant impact in the lifecycle cost of the wind farm, and hence in the cost of energy. A preliminary design of a

3-chorded jacket has been carried out, with low weight and appropriate strength. With this preliminary design as an input, time domain simulations have been produced with the specialised software BLADED. A fully integrated fatigue assessment method has been developed, consisting of calculating hot spot stress time histories from time-domain simulations of an integrated model of the complete structural set of turbine-tower-jacket, subjected to wind and waves simultaneously. First post-processing and interpretation of the simulated data show significant advantages in terms of accuracy in the fatigue assessment, by not only avoiding overdesign where the DEL method is conservative, but also by showing existing high fatigue damage where it is non-conservative.

number of cycles


hssth

1 10 100 1000 10000 100000 1e+006 1e+007 1e+008 1e+009

100

10

1

0.1

0.01

str

ess

ran

ge

bin

[M

Pa]

ACKNOWLEDGEMENTSThe authors would like to thank GL Garrad Hassan for their valuable inputs to this project.


REFERENCES [1] European Wind Energy Association, “Wind in our sails: the coming of Europe’s offshore wind energy industry,” European Wind Energy Association, Brussels, 2011.

[2] M. Perry and H. Krisdani, “Foundation Structures for Offshore Wind Turbines,” Keppel Offshore & Marine Technology Centre, Singapore, 2010.

[3] A. B. Andersen, “Why larger wind turbines improves cost of energy for offshore wind,” in Proceedings of EWEA Offshore 2011 Conference, Brussels, 2011.

[4] W. Dong, T. Moan and Z. Gao, “Long term fatigue analysis of multi-planar tubular joints for jacket-type offshore wind in time domain,” Engineering Structures, 33:6 (2011) pp. 2002-2014.

[5] Det Norske Veritas, “Offshore Standard DNV-OS-J101: Design of Offshore Wind Turbine Structures,” Det Norske Veritas, Norway, 2011.

[6] International Electrotechnical Commission, “International Standard IEC 61400-3 Wind Turbines - Design requirements for offshore wind turbines,” International Electrotechnical Commission, Geneva, 2009.

[7] M. Seidel and F. Ostermann, “Validation of Offshore load simulations using measurement data from the DOWNVInD project,” in Conference Proceedings European Offshore Wind, Stockholm, 2009.

[8] Garrad Hassan & Partners Ltd., “Bladed Theory Manual (version 4.0),” Garrad Hassan & Partners Ltd., Bristol, 2010.

[9] M. Seidel, M. von Mutius, P. Rix and D. Steudel, “Integrated analysis of wind and wave loading for complex support structures of Offshore Wind Turbines,” in Offshore Wind, Copenhagen, 2005.

[10] Det Norske Veritas, “Recommended Practice DNV-RP-203 - Fatigue design of offshore steel structures,” Det Norske Veritas, Norway, 2011.

[11] J. Jonkman, S. Butterfield, W. Musial and G. Scott, “Definition of a 5-MW Reference Wind Turbine for Offshore System Development,” National Renewable Energy Laboratory, Colorado, US, 2009.


78 KOMtech Technology Review 2012 featuRed aRticles


Comparison of SNAME T&RB 5-5A Framework and a Plasticiy-based Spudcan Model for Jackup Foundation Assessments 79

the guidelines foR site-specific assessment of moBile JacKup units given in SNAME T&RB 5-5A provide a detailed framework for modelling soil-structure interactions and performing foundation stability checks. The SNAME method has several levels with an increasing degree of sophistication. The recommended method has been used in practice and deemed robust and consistent for routine assessment purposes. While the SNAME method has gained acceptance in the industry, advanced plasticity models that describe the entire load-displacement behaviour of the spudcan footing have also evolved and become more established. However, the latter is still rarely used in practice due to the perceived complexity involved in their implementation and integration with the jackup structural model. This paper compares the use of a plasticity model with the SNAME method for jackup foundation assessments. Some sample applications in clay and sand are demonstrated with discussion on the resulting foundation responses and their implications to the foundation acceptance.

Comparison of SNAME T&RB 5-5A Framework and a Plasticity-based Spudcan Model for Jackup Foundation Assessments

okky puRwana, PhD, M.Eng, B.Eng Keppel Offshore & Marine Technology Centre michael peRRy, PhD, C.Eng Keppel Offshore & Marine Technology Centre matthew quah, PhD, C.Eng, CMarEng, FIMarEST Keppel Offshore & Marine Technology Centre

mark cassidy, DPhil, BE, FIEAust, FTSE Centre for Offshore Foundation Systems, The university of Western Australia, Australia

presented at the 13th international conference: the Jackup platform 13 - 14 september 2011, city university london, ulc.


INTRODUCTION SNAME T&RB 5-5A recommended practice [1] is routinely used for site-specific assessment of jackups. A framework to incorporate soil-structure interaction effects, which in most circumstances benefit jackup stability, has been provided in sufficient details for practical applications. Simplifications made in the foundation analysis method, particularly for dealing with non-linear foundation fixity, enable the framework to be integrated with structural models for global response analysis in a relatively straightforward manner. The fact that no empirical parameters have to be defined by the users has also allowed consistent analysis results.

In situations where the foundation still does not meet the acceptance criteria with the “standard” assessment method provided, SNAME allows use of more advanced foundation analyses such as full non-linear or load-displacement models. Recent developments for implementing the load-displacement behaviour of spudcans have concentrated on establishing numerical models with a plasticity framework. Incorporation of these models into a structural analysis code has enabled assessments of the ultimate foundation capacity of a jackup under extreme environmental loading. Despite its robustness, implementation in routine jackup assessments is still very limited due to the perceived complexity involved. As an alternative to more sophisticated analysis, if the foundation capacity envelope is exceeded it is common practice to expand the allowable bearing capacity at the expense of foundation settlement. This “displacement check” results in additional foundation penetrations and corresponding additional loads to be accommodated by the structure.

Interestingly, there is no study currently available in the public domain comparing the current industry

practice and any advanced load-displacement model. It is believed that such a comparison will provide an insight of any inherent safety margin in the recommended practice. In this study, the methods are compared for two example cases, one in sand and the other in clay with deep penetration.

OVERVIEW SNAME Foundation Analysis Framework Three levels of foundation stability assessment method are provided in SNAME. Each method has increasing complexities and can be applied in order. When a check does not meet the acceptance criteria, a higher level of check can be used. This involves more sophisticated treatment of the foundation.

By assuming the spudcan behaves as a pinned footing within the analysis, Step 1 check may be used as the simplest assessment method comprising preload capacity (Step 1a) and sliding (Step 1b) checks for leeward and windward legs respectively. In Step 2a, still with pinned footing analysis, the foundation capacity check is introduced to replace the preloading check. Where a global analysis package including suitable soil-structure interaction is not available or soil fixity cannot be relied on, Step 2a check is normally performed. The foundation capacity is essentially a combined vertical and horizontal bearing capacity envelope which can be derived from classical bearing capacity solutions.

However, it is a common practice to skip the Step 2a and directly use the Step 2b check whenever the foundation is able to provide a substantial amount of initial soil fixity. This approach includes linear vertical and horizontal springs in addition to non-linear rotational springs, where the rotational stiffness is reduced as a function of the vertical, horizontal and moment reactions. In the global response analysis, iteration is typically required to

Preload base reaction is not necessasily a representation of the ultimate capacity or failure load of the foundation. When the preload base reaction is exceeded, extra soil capacity may be available at the expense of additional footing penetration.

ensure compliance of the resulting footing reaction with the yield surface and reduced rotational stiffness function. The criteria of foundation acceptance are bearing and sliding capacity utilizations being less than unity (with respect to allowable foundation capacity). In this case, the same foundation and sliding capacity envelopes as in Step 2a are used. The allowable (factored) capacity is constructed by scaling down the initial “ultimate” capacity (unfactored) with respect to a selected origin with the use of appropriate safety factors provided in SNAME.

The method for handling Step 2b as recommended in SNAME is essentially a simplification of a full load-displacement plasticity framework to enable straightforward integration into the structural model. Since the leg–hull connection moment is typically the most critical element affecting the jackup structural integrity in the elevated condition and is heavily influenced by the foundation rotational stiffness, special attention is given to reasonably capture the moment softening with increasing load. In this regard, the vertical and horizontal displacement is assumed elastic. The use of linear translational stiffness also implies that any plastic deformations (that can potentially lead to spudcan sliding or penetration failure) are not directly captured. The use of a single yield surface also tends to make the moment reaction reduce to zero at a faster rate. This results in higher vertical and horizontal reactions.

Whenever Step 2b check is not satisfied, SNAME allows use of any justifiable more advanced method which best described the spudcan behaviour. Such methods may incorporate non-linear translational and rotational stiffnesses or a fully non-linear load-displacement model. Unlike for the lower levels, SNAME does not provide a detailed assessment method for performing Step 3. In practice, where necessary, Step 3 is often implemented by expanding the foundation capacity envelope in proportion to the preload base reaction, such that the maximum footing reaction is encompassed within the new factored envelope. The associated footing settlement resulting from the required “virtual” preload base reaction is then determined from the spudcan penetration curve by reading off the change in penetration depth due to an increase in the effective preload base reaction. The footing settlement is then checked against the structure’s tolerable displacement. This

practical approach essentially assumes vertical plastic settlement to occur when the footing reaction goes beyond the factored foundation capacity envelope.

According to SNAME, vertical bearing capacity of the foundation is defined as proven foundation reaction achieved during preloading. The preload base reaction (VLo) determines the size of yield surface and combined bearing capacity envelope. The yield surface and foundation capacity envelope are constructed from different formulations which results in somewhat different shape. The yield surface, defined in vertical (V), horizontal (H) and moment (M) VHM load space, is used for fixity iteration in the global response analysis while the foundation capacity envelope is involved for assessing the resulting final VH footing reactions.

In view of the above philosophy, preload base reaction is not necessarily a representation of the ultimate capacity or failure load of the foundation. When the preload base reaction is exceeded, extra soil capacity may be available at the expense of additional footing penetration. In situations where there is no punch-through potential or the level of settlement associated with excessive reaction is relatively small, the effective bearing capacity can increase significantly and sliding of windward footing often becomes the governing factor for the foundation stability.

Plasticity Model for Spudcan FoundationIn the strain-hardening plasticity model, any plastic penetration of the spudcan into the soil establishes a yield surface in combined loading space. The shape of the yield surface remains the same while the size increases as the foundation is pushed further into the soil. Preloading of the foundation creates a size of yield surface proportional to the achieved preload base reaction. Originating from the still water reaction, any loading within the yield surface is assumed elastic while the loading path that intersects (and expands) the initial yield locus will result in plastic deformation. During elasto-plastic loading, the reaction point must remain on the expanding yield surface. This requires the four major components of a plasticity-based model to be satisfied numerically namely, yield surface, elastic deformation, hardening law, and flow rule. The plastic displacement itself is determined from consistency between the yield surface shape, the flow rule and the hardening law.



Details of plasticity based models and development of numerical models for jackup applications can be found in numerous papers (including [2]-[8], etc) and will not be discussed here.

In its implementation for jackup foundations, the plasticity model is used as a “macro model” to define the load-displacement response of each spudcan. In the extreme loading events, the non-linear nature of foundation response, in terms of reactions and deformations, is captured as the load applied to the jackup is applied incrementally. This is in contrast to the simplified method where the plastic displacement is assumed to occur upon the exceedance of allowable foundation capacity.

CASE STUDYA case study is conducted in the present study to compare the foundation assessment method recommended in SNAME and a load-displacement plasticity model for assessing jackup foundation stability. Examples are conducted for both the sand and the clay case. The example cases in this study were selected as they do not satisfy the Step 2b check. Therefore, a higher level check, or “displacement check”, is required.

The SNAME foundation analysis framework is implemented in the computer program JUSAFE [9]

while SOS3D [10] was used for the Step 3 check using the load-displacement plasticity model. SOS3D uses the force-resultant model called ISIS which was developed from strain-hardening plasticity theory [7, 8]. Model “B” [4, 5] and Model “C” [6] are available in ISIS and represent the spudcan macro element in clay and sand respectively. In the discussion hereinafter, the term “SNAME method” is used when referring to the SNAME’s assessment technique using Step 2b and where needed a displacement check, while the SOS3D force-resultant model is termed “plasticity model”.

Analysis modelTo demonstrate the footing behaviour under storm loading, a global response analysis is carried out with a generic model jackup supported by either a clay or sand foundation. The jackup model used in the analysis has a leg-to-leg spacing of 39.3m (longitudinal) and 43.4m (transverse). In both computer programs, the jackup is represented by a

bar-stool model consisting of a planar triangular hull supported by three equivalent stick legs. The leg to hull connection is set to be rigid for simplicity. No structural assessment is made in this exercise. The environmental load is applied as quasi-static push over load and is increased in a stepwise manner while the overall stiffness is iterated to achieve the system equilibrium. In the SNAME method, this is essentially to ensure the footing reaction in the combined VHM load space complies with the yield surface and the rotational stiffness reduction function. In the plasticity model, consistency with the plasticity model as well as overall equilibrium must be maintained, as briefly described in the previous section. The elevated configuration of the jackup model for the sand and clay case is tabulated in Table 1 and illustrated in Figure 1.

The model spudcan is 15.2m in diameter with a sloping base and centre tip. Medium dense sand with friction angle of 30 degree and effective unit weight of 9.0 kN/m3 is chosen to represent a sandy site. In the sand case, a full contact of the spudcan base is made possible by applying a suitable preload level. For the clay case, normally consolidated clay with shear strength profile of 5 + 2z kPa (where z is depth in meter) and effective unit weight of 7.0 kN/m3 is adopted resulting in deep leg penetration under the same preload level applied to the sand case. The measured spudcan tip penetration is 2.9m and 25.8m for the sand and clay case, respectively.

Despite SOS3D featuring options for advanced model parameters to suit experimental data (see [5, 6] for details), the yield surface shape is set in these examples to be nearly identical to SNAME’s yield surface (as used in JUSAFE and the SNAME method). This provides a more consistent comparison and identification of the effect of the full non-linear stiffnesses on footing deformation which is not captured explicitly in the SNAME method. The “ultimate” foundation capacity parameters (VLo, Ho, Mo) and initial stiffnesses are derived from SNAME formulations and applied in both models. The hardening in the plasticity model is also calibrated to make it consistent with the spudcan penetration curve of sand and clay computed with SNAME recommended bearing capacity formula.

table 1. Jackup configuration selected in the study

sand Clay

Water depth 106.7 m 85.3 m

Airgap + Leg holding system distance from keel 29.7 m 29.7 m

Spudcan tip penetration 2.9 m 25.6 m

Effective leg length 137.9 m 137.9 m

Still water reaction 42.7 MN 42.7 MN

Preload base reaction 74.8 MN 74.8 MN

Initial foundation stiffness Kv (MN/m) 1.5E+03 3.4E+03 Kh (MN/m) 1.3E+03 2.5E+03 Kr (MN.m/rad) 4.6E+04 1.7E+05


Figure 1. Jackup elevated configurations and selected storm heading

Clay casesand case

storm heading 60 deg1 Leeward - 2 Windward

storm heading 120 deg2 Leeward - 1 Windward

85.3m.

29.7m.

25.6m.39.3m.2.90m.

106.7m.

29.7m.


Loading ConditionTo ensure consistent environmental loads particularly wave/current and inertia loads being applied in the two models, resultant point loads are applied to nodes in the jackup model rather than allowing each program to run hydrodynamic loading and dynamic analysis independently. The applied loads are derived from a set of environmental condition (omni directional) representing 50-year return period storm and consisting of wind, wave, current, and inertia forces. The environmental load factor of 1.15 as suggested by SNAME has been implicitly included in the applied load. The P-delta effect is inherently accounted for in the programs. The jackup elevated configuration is then defined in such a way that the applied load results in the maximum footing reaction at the most critical direction marginally passing Step 2b. This serves as the base case in which the level of applied load corresponds to a “load multiplier” (LM) of 1.0. In both soil cases, the loading condition is of moment-dominating force considering the utilization of leg length as well as the elevation and magnitude of applied loads.

A higher load multiplier is subsequently applied by scaling up all the components of environmental load until foundation system “failure” according to SNAME criteria is observed. The analysis results for the most critical directions for foundation stability, i.e. 60 degree from bow (resulting in 1-leeward and 2-windward legs), and 120 degree from bow (2-leeward and 1-windward legs) are highlighted here (see Figure 1). For the sand case considered in this study, after expansion of the foundation capacity in the displacement check, sliding of the windward leg at 120 degree storm heading is found to be the governing factor. On the other hand, an additional penetration of some 1.5m at the leeward leg, arbitrarily assumed to be the maximum tolerable displacement of leg in storm survival condition, is the limiting factor for the clay case under 60 degree storm heading.

Model CalibrationConsistency in the structural response between the two models are verified by setting the footings to a pinned condition. The footing response, in terms of vertical and horizontal reactions, and the structural stiffness reflected by the hull sway, under quasi-static push over load are compared. The load is the same

set of environment load to be applied for the analysis with soil fixity. As can be seen in Figure 2, under the pinned footing condition, both models show identical hull sway indicating comparable stiffness of the model jackup used therein. Identical footing reactions (not shown) are also observed.

Discussion: Sand caseThe footing load path in VH plane for the sand case with 60 degree and 120 degree storm heading is shown in Figures 3 and 4 respectively. In the figures, the load path is plotted along with the foundation capacity envelope and the yield surface (M=0). With a load multiplier of 1.0, the maximum footing reaction of the leeward leg under 60 degree storm heading lies on the factored capacity envelope and marginally satisfies Step 2b check. Further loading results in the onset of additional penetration according to the simplified Step 3 check. Given the low steepness of the spudcan penetration curve of the sand case, the calculated additional penetration of the leeward leg footing upon exceeding the initial (factored) foundation capacity is still well within the assumed structural tolerance even when a somewhat higher load multiplier is applied. The leeward footing reaction exceeds the unfactored envelopes at LM=1.36. Storm analysis must however be conducted for other storm headings and therefore sliding of the windward leg at different storm heading may be the limiting factor when a higher load is applied. At 120 degree storm heading, the windward leg reached its allowable sliding capacity at LM=1.72 while the overturning stability check is still satisfied. This is deemed the ultimate condition for the foundation stability for this sand case.

Figure 2. Model calibration

Applied total horizontal load (Mn)

3.0

2.5

0

1.5

1.0

0.5

0.0

hu

ll s

way

(m)

0.0 2.0 4.0 3.0 8.0 10.0 12.0

pinned footing

Jusafe

sos_3d

More detailed behavior of footing response in terms of reaction and deformation are shown in Figures 5 and 6 and plotted against load multiplier. The vertical footing reactions obtained from the two models are in good agreement despite the plasticity model showing a linear portion until the foundation first yields. A more substantial difference is observed in the horizontal footing reaction. After the first yield, non-linearity of the response and footing reaction redistribution is more prominent in the plasticity model than when using the SNAME method. This can be attributed to the use of linear horizontal stiffness in the latter. This results in horizontal reactions in the SNAME method that are somewhat higher at the leeward leg while lower at the windward leg (compared to the reactions obtained from the plasticity model). The sliding failure of windward leg, however, is typically more sensitive to vertical than horizontal reaction.

In terms of moment reaction, both models suggest the same response for the initial elastic portion. While the plasticity model remains elastic until the foundation first yields, the SNAME method shows moment softening behavior earlier. After reaching the peak sharply and dropping, the plasticity model produces fairly similar post-yield response to that of the SNAME method at the windward legs. At the leeward legs, however, as the vertical load increases the SNAME model’s moment reaction eventually drops to zero whilst the plasticity

model is able to maintain some capacity. The latter is contributed by the yield surface expansion as the plastic vertical deformation takes place at the leeward leg. On the other hand, adoption of a single fixed-sized yield locus in SNAME results in the moment reaction at the leeward leg being forced to drop as the vertical reaction increases to balance the applied load. Exceedance of the yield surface will start immediately after the moment reaction falls to zero and the footing effectively becomes pinned.

Unlike in the SNAME method whereby the translational deformation follows a linear relationship, the plasticity model suggests an increasing rate of displacement after the first yield. At LM=1.72 whereby the foundation stability is considered at its limit, the vertical displacement of the leeward leg predicted by the plasticity model is smaller than the additional penetration derived from the simplified Step 3 check. The latter is derived from the spudcan penetration associated with “virtual” preload reaction in order to encompass the maximum footing reaction within the new factored envelope. In this case, a 30% increase in the preload base reaction is required which translates to an additional penetration of some 0.5m computed from the spudcan penetration curve. This suggests that for the sand case considered here the SNAME method gives a conservative prediction of footing settlement compared to the plasticity model results.


Figure 3. Footing reaction for sand case under 60 degree storm heading

100

90

80

70

60

50

40

30

20

10

0

V (M

n)

0 2 4 6 8 10 12h (Mn)

Stbd

Fwd

Factored Capacity

Unfactored Capacity

Factored Sliding

Unfactored Sliding

Yield Surface

PortLM = 1.72

LM = 1.36

LM = 1.00

Jusafe

sos_3d

Fwd

Fwd

Port

Port

Stbd

Stbd

Figure 4. Footing reaction for sand case under 120 degree storm heading

100

90

80

70

60

50

40

30

20

10

0

V (M

n)

0 2 4 6 8 10 12

Jusafe

sos_3d

Fwd

Fwd

Port

Port

Stbd

Stbd

Factored Sliding

Unfactored Capacity

Unfactored Sliding

Factored Capacity

Yield Surface

Stbd

FwdPort

LM = 1.00

LM = 1.36

LM = 1.72

h (Mn)


Figure 5. load and deformation for sand case under 60 degree storm heading

Fwd

Port

90

80

70

60

50

40

30

20

10

V (M

n)

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

Stbd

LM = 1.00

LM = 1.36

LM = 1.72

load Multiplier

0.15

0.1

0.05

0

-0.05

Vert

ical

dis

pla

cmen

t, d

w (m

)

LM = 1.72

LM = 1.36

LM = 1.00 Port

Fwd

Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

0.15

0.1

0.05

0

-0.05

Vert

ical

dis

pla

cmen

t, d

w (m

)

LM = 1.72

LM = 1.36

LM = 1.00 Port

Fwd

Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

0.15

0.1

0.05

0

LM = 1.72

LM = 1.36

LM = 1.00 Port

Fwd

Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

ho

rizo

nta

l dis

pla

cmen

t, d

w u

(m)

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

5

4

3

2

1

0

h (M

n)

LM = 1.72

LM = 1.36

LM = 1.00

Port

Fwd

Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

2

1.5

1

0.5

0

ro

tati

on

, °(d

eg)

Port

Fwd

Stbd

LM = 1.72

LM = 1.36

LM = 1.00

LM = 1.72

LM = 1.36

LM = 1.00

Fwd

Stbd

Port

load Multiplier

100

90

80

70

60

50

40

30

20

10

0

M (M

n.m

)

Jusafe Fwd Port Stbd sos_3d Fwd Port Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

Figure 6. load and deformation for sand case under 120 degree storm heading


80

70

60

50

40

30

20

10

0

V (M

n)

h (M

n)

100

90

80

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60

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40

30

20

10

0

M (M

n.m

)

5

4

3

2

1

0

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

LM = 1.72

LM = 1.36

LM = 1.00

Port

Fwd

Stbd

LM = 1.72

LM = 1.36

LM = 1.00Port

Fwd

Stbd

LM = 1.72

LM = 1.36

LM = 1.00

Port

Fwd

Stbd

LM = 1.72

LM = 1.36

LM = 1.00

Port

Fwd

Stbd

LM = 1.72

LM = 1.00

Port

Fwd

Stbd

LM = 1.36

LM = 1.72

LM = 1.36

LM = 1.00

Port

Fwd

Stbd

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier load Multiplier

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00

load Multiplier

0.15

0.1

0.05

0

-0.05

Vert

ical

dis

pla

cmen

t, d

w (m

)

0.15

0.1

0.05

0

ho

rizo

nta

l dis

pla

cmen

t, d

w u

(m)

2

1.5

1

0.5

0

ro

tati

on

, °(d

eg)


0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00


The horizontal displacement is significantly lower for the SNAME method due to the elastic nature of horizontal stiffness adopted. Despite the underestimation, the effect of actual footing lateral displacement on the structure may not be critical as SNAME also provides for leg eccentricity equivalent to 0.5% of the leg length below the hull. In this case, this leg eccentricity of more than 0.6m, is far greater than the 0.1m predicted by the plasticity model at the sliding limit of LM=1.72.

The footing rotations predicted by both models agree very well until the reaction in the SNAME method reaches the foundation capacity envelope. When loaded further, beyond the initial capacity, the SNAME method produces a constant rate of increase in footing rotation as the rotational stiffness is reduced to a pinned footing case. In consideration of the simplification made in the SNAME framework the moment-rotation behavior is predicted reasonably well and tends to be on the conservative side compared to the plasticity results. This also implies that the SNAME method is able to model the influence of soil fixity on the leg-hull connection moment in a reasonable way.

It is worth noting that generally in sand significant reserve capacity at the leeward footing maybe available upon the exceedance of factored yield surface and that the associated settlement is relatively small. In this situation, the sliding of the windward footing often becomes the critical check for the foundation stability.

Discussion: Clay caseFigure 7 shows the footing load path of the clay case for 60 degree storm heading. Unlike the sand case, whereby the sliding of windward leg governs, settlement of the leeward leg becomes the limiting factor for the clay case and therefore the following discussion is focused solely on the 60 degree storm heading.

Upon the footing reaction reaching the factored envelope at LM =1.0, the jackup is loaded further until the leeward footing becomes pinned or literally the footing reaction lies on the yield surface on V-H plane at M=0, at LM=1.18. At this point, between LM=1.0 and LM=1.18 loading state, an excessive

additional penetration of the leeward footing has occurred according to the simplified Step 3 check. The calculated “virtual” preload base reaction suggest that an additional penetration of 2.5m will have occurred at LM = 1.18. To limit the settlement within the structural limit of 1.5m, the jackup can only withstand 10% extra load (LM = 1.10).

As illustrated in Figure 8, similar to the sand case fairly comparable vertical reaction response and discrepancy of horizontal reaction are observed from the two models. Both models, however, produce more consistent moment response at all the three legs than in the sand case. Despite the plasticity model reaching higher peak moment at the leeward leg, the post-yield behavior of all the legs show very similar paths until the leeward footing goes to pinned condition.

In terms of footing deformation, though the plasticity model predicts much lower footing settlement than the simplified Step 3 check upon exceedance of the factored yield surface at LM = 1.18, it however indicates the onset of apparent “plunging” failure of the leeward leg. This can be observed from the vertical and horizontal displacements which increase exponentially at this level of load. In this case, the inherent conservatism of SNAME method in grossly predicting additional penetration is in a way able to prevent the apparent failure of the leeward footing captured by the plasticity model.

Figure 7. Footing reaction for clay case under 60 degree storm heading

100

90

80

70

60

50

40

30

20

10

0

V (M

n)

0 2 4 6 8 10 12 14 16 18h (Mn)

Stbd

Fwd

Factored Capacity

Unfactored Capacity

Factored Sliding

Unfactored Sliding

Yield Surface

Port

LM = 1.18

LM = 1.00

Jusafe

sos_3d

Fwd

Fwd

Port

Port

Stbd

Stbd


Figure 8. load and deformation for clay case under 60 degree storm heading

90

80

70

60

50

40

30

20

10

V (M

n)

h (M

n)

160

140

120

100

80

60

40

20

0

M (M

n.m

)

5

4

3

2

1

0

load Multiplier


load Multiplier

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4


0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4


0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

0.08

0.06

0.04

0.02

0

-0.02

0.08

0.06

0.04

0.02

0

1.4

1.2

1

0.8

0.6

0.4

0.2

0

ro

tati

on

, °(d

eg)

Vert

ical

dis

pla

cmen

t, d

w (m

)h

ori

zon

tal d

isp

lacm

ent,

dw

u (m

)

LM = 1.18

LM = 1.00 Port

Fwd

Stbd

LM = 1.10

LM = 1.18

LM = 1.00 Port

Fwd

Stbd

LM = 1.10

LM = 1.18

LM = 1.00

Port

Fwd

Stbd

LM = 1.10

LM = 1.18

LM = 1.00 Port

Fwd

Stbd

LM = 1.10

LM = 1.18

LM = 1.00

Port

Fwd

Stbd

LM = 1.10

LM = 1.18

LM = 1.00

Port

Fwd

Stbd

LM = 1.10


SUMMARYA comparison between the SNAME method and the plasticity model for assessing foundation stability of jackup under storm condition in sand and clay foundation has been performed. Although the two methods handle the non-linear behaviour of the soil response in a different way, in general the SNAME foundation assessment framework is able to produce assessment results which compare well with those obtained using the plasticity model.

The practical Step 3 approach, i.e. expansion of foundation capacity envelope through the “virtual preload” concept to estimate footing settlement, is found conservative compared to the plasticity approach. On the other hand, as expected the linear horizontal stiffness tends to under predict horizontal displacement. However, the additional horizontal displacement predicted by the plasticity model may not be large compared to the allowance made for initial leg inclination.

Although the two methods handle the non-linear behaviour of the soil response in a different way, in general the SNAME foundation assessment framework is able to produce assessment results which compare well with those obtained using the plasticity model.

For the sand case, when loaded beyond the factored envelope significant extra foundation capacity is available. The foundation reaches its ultimate stability when sliding of the windward footing occurs rather than excessive additional settlement or “plunging” failure of the leeward footing. In the clay case, an overestimation of settlement level at the leeward footing is observed from the simplified Step 3 check as compared to the plasticity model. However, at the same time this conservatism is also able to prevent the onset of foundation failure at the leeward leg, as captured by the plasticity model.

The findings observed in this study are based solely on deep water cases and the moment-dominating force situation. Further studies with other possible jackup or loading configurations such as shallow water cases with high base shear are necessary to verify the footing behaviour captured in this exercise.

ACKNOWLEDGEMENTSThe authors would like to thank Dr Henry Krisdani of KOMtech for his assistance with the analytical tool for performing foundation analysis and check. The computer program SOS3D was written in collaboration between the fourth author and Associate Professor Britta Bienen of The university of Western Australia. It also uses the ISIS soil models, originally coded by Professor Guy Houlsby of the university of Oxford together with the fourth author. The fourth author acknowledges the support of the Australian Research Council through their Future Fellowship and Centres of Excellence Programmes and The Lloyd’s Register Educational Trust. Permission from Keppel Offshore & Marine for publishing the paper is gratefully acknowledged. The author would also like to acknowledge the organizer of the 13th International Conference: The Jackup Platform (City university London, uK) for the permission to reproduce this paper.


REFERENCES [1] SNAME (2008). Guidelines for site specific assessment of mobile jack-up units. Society of Naval Architects and Marine Engineers, Technical and Research Bulletin 5-5A Rev. 3, New Jersey.

[2] Schotman, G.J.M. (1989). The effects of displacements on the stability of jackup spudcan foundations. Proc. 21st Offshore Technology Conf., Houston, OTC 6026.

[3] Van Langen, H., Wong, P.C., Dean, E.T.R. (1999). Formation and validation of a theoretical model for jack-up foundation load-displacement analysis. Marine Structures, Vol. 12(4), pp. 215-230.

[4] Martin, C.M. and Houlsby, G.T. (1999). Jackup units on clay: structural analysis with realistic modelling of spudcan behaviour. Proc. 31st Offshore Technology Conf., Houston, OTC 10996.

[5] Martin, C.M. and Houlsby, G.T. (2001). Combined loading of spudcan foundations on clay: numerical modelling. Géotechnique 51(8): 687-700.

[6] Houlsby, G.T. and Cassidy, M.J. (2002). A plasticity model for the behaviour of footings on sand under combined loading. Géotechnique, 52(2): 117-129.

[7] Houlsby GT. (2003). Modelling of shallow foundations for offshore structures. In: Proceedings of the international conference on foundations, Dundee, Scotland.

[8] Cassidy M.J., Martin, C.M., Houlsby, G.T. (2004). Development and application of force resultant models describing jack-up foundation behaviour. Marine Structures, Vol. 17, pp. 165–93.

[9] Perry, M.J. (2010). JUSAFE R4.1 – Jack Up Site Assessment with Fixity Evaluation software. User Manual. Keppel Offshore & Marine Technology Centre.

[10] Bienen, B. and Cassidy, M.J. (2006). Advances in the three-dimensional fluid-structure-soil interaction analysis of offshore jack-up structures. Marine Structures, Vol. 19 (2-3), pp. 110-140.



Use of Field Penetrometer Data for an Integrated Jackup Installation System 93

JacKup geotechnical hazaRds such as unpRedicted leg penetRation, rapid leg penetration and punch-through continue to occur at an increasing rate despite efforts to minimise these risks. Hence, it is essential to improve installation guidelines and site specific assessment to achieve safe jackup rig installation. Assessment of spudcan penetration is one of the key aspects required in a jackup site specific assessment. An accurate spudcan penetration prediction underpins reliable site specific assessment. In this paper, we are reviewing the current practice for spudcan penetration prediction. Design approaches to obtain spudcan penetration curve from field penetrometer data are discussed. The spudcan penetration prediction design approaches are then incorporated into an integrated jackup installation system. The system calculates the spudcan penetration curve based on penetrometer data and then closely monitors progression of spudcan installation. The aim is to assist jackup operators in making decision and taking measures to prevent or mitigate potential geotechnical hazards, particularly punch-through.

Use of Field Penetrometer Data for an Integrated Jackup Installation System

henry KRisdani, PhD, B.Eng Keppel Offshore & Marine Technology Centre

okky ahmad puRwana, PhD, M.Eng, B.Eng Keppel Offshore & Marine Technology Centre matthew quah, PhD, CEng, CMarEng, FIMarEST Keppel Offshore & Marine Technology Centre shazzad hossain, PhD, M.Eng Centre for Offshore Foundation Systems, The university of Western Australia, Australia mark Randolph, PhD, MA, FAA, FTSE, FREng, FIEAust, CPEng Centre for Offshore Foundation Systems, The university of Western Australia, Australia

presented at the 13th international conference: the Jackup platform. 13 - 14 sep 2011, city university london, uK.



INTRODUCTION Most of the world’s offshore drilling in shallow water (up to 150m) is performed from self-elevated mobile jackup rigs. A unique feature of jackup rigs is their self-installing capability which differentiates the design of its spudcan foundations from conventional onshore foundations. Prior to commencing jackup operations, spudcans are routinely proof loaded by static vertical preloading (either sequentially or simultaneously) to increase the size of the yield envelope in vertical, moment, horizontal load space. This helps to ensure they have sufficient reserve capacity in any design extreme storm event (SNAME [1]). Typically preloading is accomplished by pumping seawater into holding tanks within the hull. This causes the spudcan foundations to penetrate into the seabed until the load on the spudcan is equilibrated by the resistance of the underlying soil. The preload is then dumped and the hull is elevated to provide an adequate air-gap for subsequent operation.

Due to the method of installation, several potential geohazards can arise during installation of jackup foundations. Sharples et al. [2], MSL [3] and Jack et al. [4] described in detail typical geohazards and their statistics collating reported case histories. Rapid leg penetration and spudcan ‘punch-through’ were identified as the most common geohazards that may lead to serious consequences. Installing and preloading a jackup in stratified deposits, where a strong layer overlays weaker soil, remains a challenge for the offshore industry. Rapid leg penetration may lead to buckling of the leg, effectively decommissioning the platform, or may even result in toppling of the unit (Aust [5]; Maung & Ahmad [6]; Brennan et al. [7]; Kostelnik et al. [8]; Chan et al. [9]). In addition, the jackup may also collide with the adjacent platform. Punch-through may be defined as sudden leg penetration due to a drop in bearing capacity, and where the legs cannot be jacked fast enough to maintain the hull levelness. The extent to which rapid leg penetration may be

controlled, avoiding punch-through, has improved with advances in jackup technology and geotechnical knowledge.

The risk of punch-through may be mitigated by appropriate operational procedures applied by the jackup operator during installation, but the risk increases if the spudcan penetration prediction is inaccurate or if the load-penetration response is not monitored carefully. The former is associated with the quality of geotechnical survey and assessments conducted for the intended site. Inadequate site-specific geotechnical information, improper geotechnical survey, poor selection of engineering parameters from the geotechnical data, and inappropriate methods of bearing capacity prediction are several factors that may eventually result in poor leg penetration prediction. Conversely, provision of accurate penetration prediction may become ineffective if the actual load penetration response is not tracked during installation. Without the ability to monitor the response, occurrence of unpredicted spudcan behaviour cannot be observed and effective preventive measures cannot be taken in time. This paper presents a review of current practice of spudcan penetration prediction and proposes methods for determination of spudcan penetration curve from penetrometer data. An integrated system that combines spudcan penetration prediction with a monitoring system during jack-up installation is also presented.

SPUDCAN PENETRATION PREDICTION: CURRENT PRACTICE Penetration resistance of spudcan foundations is commonly assessed in the framework of bearing capacity formulations (SNAME [1]; InSafeJIP [10]; ISO [11]). This is a two-step approach in which soil strength parameters are derived from the site specific soil investigation data (comprising tests on borehole samples and in-situ penetrometer testing) for use in bearing capacity models. Spudcan foundations undergo progressive penetration during preloading,

The system calculates the spudcan penetration curve based on penetrometer data and then closely monitors progression of spudcan installation.


contrasting with onshore practice where a footing is placed at the base of a pre-excavated hole or trench. However, spudcan resistance profiles are generally assessed within the framework used for onshore foundations, considering the bearing capacity of a series of ‘wished-in-place’ spudcans at successively increasing depths with any stratification or strength anomalies unchanged from the original in-situ conditions, and with all soil within the plan area of the spudcan down to the base level conveniently removed or replaced by a surcharge of zero strength backfilled soil above the spudcan (SNAME [1]). Recently, mechanism based design approaches have been developed for single and double layer soils taking into account the evolving soil failure patterns over the full penetration depth (Randolph & Gourvenec [12]). This section reviews current practice of spudcan penetration prediction for two-layer and multi-layer soil systems which are prone to punch-through hazard.

Two-Layer SoilsAs discussed, the potential for unexpected punch-through failure of a jackup exists during installation and preloading in layered soils. Soil conditions of a thin layer of stiff clay or sand overlying a weaker stratum of clay are particularly hazardous. The former stratification is common in the Sunda Shelf, offshore Malaysia and offshore Thailand and the latter in the North Sea, offshore India, offshore Australia and in the Arabian Gulf.

Stiff-over-Soft ClayWished-in-place approach: For assessing spudcan penetration response on stiff-over-soft clay, SNAME [1] recommends calculation of the bearing capacity, Qv, using Brown & Meyerhof ’s [13] factor, but adjusted for embedment depth by applying a constant depth factor, following the semi-empirical approach of Skempton [14]. Brown & Meyerhof ’s approach was developed on the basis of model tests conducted on a surface circular footing. Based on the results from small strain FE analysis of a surface

circular footing, Edwards & Potts [15] also suggest a design approach. These approaches do not account for the distortion of the upper layer as it punches through into the lower layer.

Mechanism based approach: Hossain & Randolph [16, 17] report results of an extensive investigation, combining centrifuge model tests and large deformation finite element (LDFE) analysis. It was found that the failure modes assumed by the currently available recommended practices are not consistent with those observed from centrifuge tests and LDFE analyses. Severe punch-through was associated with purely vertically downward soil displacements beneath the spudcan in the upper layer, involving punching shear, with clear shear planes in the shape of a truncated cone forming in the upper layer below the spudcan. A concise approach for assessing profiles of spudcan penetration resistance through stiff-over-soft clay was reported by Hossain & Randolph [18]. The effects of strain rate and softening on the response were accounted for.

Sand-over-ClayWished-in-place approach: For assessing spudcan penetration in sand-over-clay, SNAME [1]

recommends the ‘punching shear’ approach developed by Hanna & Meyerhof [19], but assuming vertical shear planes beneath the advancing spudcan. The product of the punching shear coefficient, Ks, and tanφ is linked to the normalised shear strength of the underlying clay layer by a simple expression as Kstanφ = 3su/γs Ds for a lower bound estimation. More accurate values of Ks as a function of φ and the ratio of bearing capacity for a surface footing on the bottom clay layer to the top sand layer, qclay/qsand were presented graphically by Meyerhof & Hanna [20] and ISO [11]. The ‘projected area’ method is suggested as an alternative design approach, with rates of spread in the range of 1:5 to 1:3 (horizontal: vertical). However, the properties of the upper sand layer are ignored in this approach.

The risk of punch-through may be mitigated by appropriate operational procedures applied by the jackup operator during installation.


Mechanism based approach: The observed mechanisms in centrifuge tests are significantly different from the assumed mechanisms in the wished-in-place approaches (Craig & Chua [21,

22]; Teh et al. [23]). The mechanism consists of a truncated cone of sand being forced vertically into the underlying clay layer, which in turn changes from pure vertical movement to radial. The outer angle of the sand frustum being forced into the clay reflects the dilation of the sand. Based on the observed mechanisms, two alternative methods have been developed by Lee et al. [24] and Teh et al. [25] for predicting the bearing capacity at the surface, peak and at the interface. The stress level and dilatant response of the sand are accounted for using an iterative approach.

Multi-Layered Soils with Interbedded Strong LayersDepletion of known reserves in the traditional regions and in shallow waters results in exploration in deeper, unexplored and undeveloped environments with more complex seabed soil conditions. In emerging provinces and fields, highly layered soils are prevalent. Over 75 % of the case study data sets forming the basis for the InSafeJIP [10] involve stratified seabed profiles, with interbedded layers of clay and sand evincing strong variations in shear strength. The Sunda Shelf, offshore Malaysia, Australia’s Bass Strait and North-West Shelf, offshore India and Arabian Gulf are particularly problematic in terms of stratigraphy and soil types. Installing and preloading a jack-up rig in stratified deposits, where an interbedded strong layer overlays a weak layer, is hazardous, with the potential for severe punch-through failure.

The bearing capacity for a spudcan penetrating in three or more soil layers can be computed by integrating the squeezing and punch-through criteria for two layer systems. In the offshore industry, two approaches are used: (a) top down approach and (b) bottom up approach. SNAME [1] recommends the latter approach. Firstly the bearing capacity of a footing resting on top of the lower two layers is computed. These two layers can then be treated as one (lower) layer in a subsequent two layer system analysis involving the third (from bottom) layer. A strength parameter at the interface is updated based on the calculated bearing capacity at the surface of the bottom two layers. For the top down approach, two

to three layers may be required need to be considered together depending on the relative thickness of the layers. A layer thickness less than the spudcan tip height is generally merged with the adjacent upper or lower layer based on the layer soil property.

Centrifuge modelling and LDFE analysis of spudcan foundations penetrating through multi-layered soils with interbedded strong layers are being investigated at COFS, UWA. Test specimens are prepared covering a range of normalised layer soil properties and geometry, mostly mimicking strength profiles from reported case histories particularly those where punch-through failure occurred. Punch-through and rapid leg penetration (for strong-over-weak) and squeezing (for the reverse) have been demonstrated by the penetration resistance profiles and associated soil failure mechanisms and provide a basis for developing design calculations.

Challenges From the above discussion, it is clear that an assessment of spudcan penetration resistance using either wished-in-place or mechanism based approaches relies primarily on (for a given analytical model) the selection of two key parameters: (a) layer boundaries for the soil profile, and (b) strength parameters (su or φ) within each layer.

Geotechnical Assessment: Variability in Measured Shear StrengthPrior to jack-up deployment at a site, geophysical and geotechnical surveys are normally conducted from a dedicated vessel. The geotechnical investigation typically consists of continuous borehole sampling, sometimes combined with continuous penetrometer tests. Alternate sampling and penetrometer testing within a single borehole is also practiced on occasions. The primary goals include layer identification, soil classification and selection of strength parameters for each layer.

Laboratory tests: With the samples taken, various offshore or onshore laboratory tests are conducted from which shear strength data and other engineering parameters are derived. These tests, particularly if performed offshore, are often relatively simple, such as unconfined compression, miniature vane (or Torvane) and pocket penetrometer tests. Significant scatter of the data is typical, so that engineering


judgment is required to derive layer boundaries and design shear strength profiles. Figure 1 shows an example of the degree of variability in terms of shear strength information for a jack-up site. Variability in the shear strength measured from several types of tests, with potentially some degree of sampling disturbance, leads to fundamental uncertainty that may in turn lead to inaccuracies in the bearing capacity calculation process.

In-situ penetrometer tests: Advances in in-situ testing, using piezocone, T-bar and ball penetrometers, enables more representative soil properties to be determined. Especially for soil conditions where undisturbed sampling is challenging, field penetrometer testing is a superior approach to derive in-situ shear strengths and other parameters. In silty material, drainage characteristic parameters can also be derived from specific procedures, such as varying the rate of penetration or by from dissipation tests (Randolph & Hope [26]), with the objective of obtaining the operational cv and useful information about soil type. With these parameters, the effect of partial drainage on the penetration resistance can be quantified, as discussed later.

In most offshore design projects, in-situ penetrometer tests (e.g. CPTu, T-bar, ball) are first correlated with element test data in order to assess a suitable

Figure 1. undrained shear strength parameters from field and laboratory tests

resistance factor, so that effectively the absolute value of penetration resistance is not used directly. For instance, at present a range of Nkt factors are applied to the net cone resistance to provide upper and lower bound values of undrained shear strength. These ranges are generally plotted together with the laboratory strength data on the borehole logs with inadequate regard for the selection of the site and soil layer specific Nkt values (Osborne [27]). Appropriate consideration should be given to the selection of the applied coefficients for each soil layer, and justified within the SI report instead of using a ‘blanket’ Nkt

value for all the soil units at a specific location or region (Erbrich [28]; Osborne [27]). The value of Nkt

may vary between 8.6 and 20 (Low et al. [29]). This approach provides a means of interpolating strengths, in terms of relative values, but the resistance factor relating penetration resistance to shear strength measured in an element test is generally derived for the given site by correlation, so that the absolute value of the penetration resistance is not used directly. Design calculations, however, are usually based on classical theoretical relationships, as previously discussed. There is, therefore, an inconsistency be¬tween the interpretation of data from in-situ penetration tests, where the correlated resistance factor may not accord with theory or a design process based on theoretical solutions.

Layer Identification The identification of potential punch-through failure, the critical depth at which it occurs and its severity, the predicted final spudcan penetration under the applied preload, and the general trend of the load-penetration response at intermediate depths are all significantly affected by layer identification and selection of design parameters. In Figure 2, spudcan penetration predictions provided by three different assessors based on the shear strength information given in Figure 1 indicate a wide range of predicted penetrations. In the jack-up industry, identification of layers and soil type (and quantifying geotechnical properties of each layer) of the seabed sediments relies heavily upon laboratory test data, with the former mostly performed by visual inspection of the various soil parameter profiles. This often results in highly subjective assessments, although this can be improved through statistical analyses (InSafeJIP [10]). It may be sufficient for sediments in the Gulf of Mexico, with minimal horizontal and

dep

th (m

)

0 20 40 60 80 100 120 140 160

undrained shear strength (kPa)

Motorvane

Motorvane (r)

torvane

Pocket Pen

triaxial uu

triaxial uu (r)

CPtu, nkt=15

CPtu, nkt=20

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Penetrometer testing equipment integrated within the jackup has been proposed by Quah et al. [33] (see Figure 3). The integrated penetrometer equipment enables piezocone penetration testing to be carried out immediately below the spudcan, prior to the preloading stage, with the spudcan soft pinned on the seabed to create a sufficiently stable base. The responses from the penetrometer are processed, displayed on a real time basis, and stored for further processing. Provision of such equipment on board jack-ups will provide the jack-up operator with resources to detect unforeseen situations and to help make the right decisions for mitigating any potential problems in the course of installation. Building upon this initiative, the present study addresses the next step: a post-processing framework that will provide layer identification, soil classification, strength assessment and finally an automated design approach to obtain the spudcan penetration curve. Potential for punch-through can be identified and suitable leg installation procedures can be planned. Although designed as an automated system, providing predicted resistance profiles in real time, involvement of appropriately qualified engineers is still seen as essential. The system will be configured in a manner that allows direct input by engineers in order to adjust layering or strength parameters as the need arises and according to relevant experience.

vertical variability (Menzies & Roper [30]), but not in emerging frontier regions with more complex soil conditions. The difficulty in obtaining high-quality soil samples from these sites for laboratory determination of soil properties has also placed increasing reliance on results from in-situ testing. Furthermore, if the sampling gap is too large, layer boundaries cannot be identified properly, especially in complex multi-layered soils.

INTEGRATED PENETROMETER EQUIPMENT IN JACKUP RIGSThe most commonly adopted in situ test for offshore site investigation for jackup installation is the piezocone penetration test (CPTu). The CPTu has major advantages over traditional methods of field site investigation, such as drilling and sampling, because it is fast, repeatable, and economical. In addition, it provides near-continuous data and has a strong theoretical background. The standard cone penetrometer is cylindrical in shape, having a conical tip with a base area of 10 cm2 (diameter Dc = 35.7 mm) and 60° tip apex angle. The penetration tests are carried out at a rate of vc = 20 mm/s. It is now common that pore pressure is measured behind the cone in what is referred to as the u2 position (ISSMFE [31]; ASTM [32]). The accuracy and precision of the cone pore-pressure measurements are reliable and repeatable as loss of saturation of the pore-pressure element is not an issue for offshore testing.

Figure 2. spudcan penetration predictions and actual data measured in the field

Figure 3. CPtu integrated with a spudcan

sp

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Vertical load capacity (x1000 tonnes)

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upper bound predictions

lower bound predictions

stillwater reaction

Preload reaction

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(b) fs

(c) u2

Figure 4. Measured qc, fs, and u2 profiles

(a) qc

dep

th (m

)

Cone resistance, qc (kPa)

0 2000 4000 6000 8000 100000

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)sleeve friction, fs (kPa)

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SPUDCAN PENETRATION PREDICTION BASED ON PENETROMETER DATA Layer Identification and Soil TypeParameters and CorrectionsParameters measured during a piezocone test include (a) cone tip resistance, qc, (b) sleeve friction, fs, and (c) pore water pressure at the shoulder, u2. Figure 4 shows profiles of these three measured (raw) parameters for a case as a function of penetration depth below the mudline (all zeroed at this depth). Due to the inner geometry of the cone, the ambient pore pressures acts on the shoulder behind the cone (referred to as the unequal end-area effect). The measured cone resistance qc is therefore corrected to total cone resistance qt using the following relationship (Lunne et al. [34])

(1)

where α is the net area ratio. Although the range of α for different cone designs is 0.59 to 0.85, the area ratio for the cone should always be determined in a calibration vessel (Campanella & Robertson [35]). The total cone resistance must then be adjusted for the overburden pressure to give the net cone resistance, q

cnet, calculated as

(2)

where σv0 is the in-situ total overburden stress (obtained by integrating γ with depth, where γ is the total unit weight of the soil).

Normalised Parameters As the effective overburden stress increases with depth, cone penetration tip resistance (and in normally consolidated soils, u2 and fs as well) also tend to increase. Therefore, normalisation of measured parameters is critical for rational evaluation of soil behaviour (Wroth [36]). Normalisation of cone tip resistance is typically based on the vertical effective stress, σ'v0, as

(3)

qt = qc + u2 ( 1 - α)

qcnet = qt - σv0

Qt1 =

qcnetqt - σv0

σ'v0= σv0


Figure 5. Algorithm for analysing CPtu data for indirect or direct design approaches

CPtu

qc, fs, u2

Calculate normalised parameters

Qtn, Fr, bq, u, ic

layer identification use statistical analysis

direct CPt to spudcan penetration profile

soil type for each layer e.g. sand, silt, Clay

use classification chart

extraction of soil strength parameters

e.g. su, φ

bearing capacity use design approaches

identify potential for punch - through failure

Alternatively, the excess pore pressure may be normalised by the vertical effective stress, σv0, referred to as normalised pore pressure and given by (Schneider et al. [42])

(6)

The sleeve friction, fs, which is considered much less reliable than the pore pressure measurement for offshore conditions (Lunne & Andersen [43]), is normalised as

(7)

The normalised parameters are used for layer identification, soil classification and extracting design strength properties for each layer. The normalisation requires one additional parameter, which is the total unit weight of the soil (γ). In absence of water content data from recovered samples, this may be estimated using either CPTu data or the measured shear wave velocity (from seismic peizocone test; SCPTu) (Mayne et al. [44]).

Robertson & Wride [37] and Zhang et al. [38] updated the above normalisation incorporating a variable stress component, n, according to

(4)

where qcnet/pa is the dimensionless net cone resistance, (pa/σ'v0)n is the stress normalisation factor, pa is the atmospheric pressure in the same units as qcnet and σ'v0, and n is the stress exponent that varies with the normalised soil behaviour type index (SBTn). Note that when n = 1, Qtn = Qt1. Expressions for calculating n in relation to SBTn have been proposed by Robertson [39]. For fine-grained soils, the stress exponent is normally taken as unity, while for coarse-grained soils it will range from 0.5 to 0.9.

The normalised measured pore pressure is expressed by the pore pressure parameter, Bq (Senneset & Janbu [40]; Robertson [41]) defined as

(5) Bq = qcnet

u2 - u0

qt - σv0=

∆ u2

U =

∆ u2u2 - u0

σ'v0= σ'v0

100%Fr =

fsfs

qt - σv0= qcnet

100%

Qtn =

qcnetn

σ'v0pa

pa( ) ( )

LUse of Field Penetrometer Data for an Integrated Jackup Installation System 101

The algorithm for processing the CPTu data is shown in Figure 5. It comprises four main steps: layer identification, selection of soil type for each identified layer, extraction of the strength profile (or parameters) for each layer, and finally evaluation of the spudcan penetration profile. The profiles of normalised parameters for the data shown in Figure 4 are plotted in Figure 6.

Layer Identification It is proposed that the CPTu data provide the basis for soil profiling, as well as parameter assessment for geotechnical design. Subjective estimates, as conventionally used today, can often lead to erroneous results. However, the usefulness of the proposed approach, which relies on established soil classification charts, relies on accurate determination of each soil layer.

A statistical method has been used to identify layer boundaries from the CPTu data. Statistical analysis can be carried out using each normalised CPTu data separately (referred to as univariate statistics), with the layers then demarcated from a weighted average of the separate results. Alternatively, and preferably, normalised parameters can be used together to undertake multivariate analysis. The three normalised parameters from a CPTu exhibit a different kind of behaviour in different types of soils and therefore any method that considers all the parameters simultaneously, rather than individually, will certainly provide more accurate demarcation of layer boundaries. Figure 7 shows an example of identified layers using multivariate statistics on the CPTu data from Figure 4. It should be noted that

Figure 6. normalised data for the case presented in Figure 4

(a) Qm

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(b) Fr

dep

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CPTu profiles were not continuous, but with some gaps. Linear interpolation was carried out between stopping and resuming data points. This has some effect on layer identification. If only the layers whose characteristics are highly dissimilar are required only the highest peaks of the statistic need to be chosen. However, if a more elaborate layer identification is necessary, even values giving moderately high values should be selected, which is unnecessary for a 10~20 m diameter spudcan penetration.

Soil ClassificationOne of the major applications of the CPTu has been the identification of soil type, for which soil behavioural classification charts using normalised piezocone parameters are a well established basis. Several charts have been proposed for evaluating soil type from CPTu data, notably Robertson [39, 41], Jefferies & Been [45] and Schneider et al. [42]. Hossain et al. [46] also reported an improved chart developed focusing particularly on spudcan foundations. A normalised soil behaviour type (SBTn) index Ic (Robertson & Wride [37]; Jefferies & Been [45]) provides a quantitative measure of the soil type, which may then be used in various correlations. For identifying the soil type for each layer, the normalised data are plotted in the classification charts, as illustrated in Figure 8 for the case presented in Figure 7. The

deduced layers and corresponding soil types are summarised in Table 1, and compared with layer boundaries and soil types from the reported borehole log (from onshore and offshore lab testing on cored soil samples). Reasonable agreement between the results from the developed framework and borehole logs from ten case histories, in the context of both layer identification and soil classification, has confirmed the accuracy and hence applicability of the procedures (Hossain et al. [46]).

Design Approaches Identified layers and soil type for each layer can be used in either an indirect (two-step) design approach, where soil strength parameters are extracted for each layer and then used for resistance calculations, or in a more direct design approach based on the cone resistance data itself.

Indirect Design ApproachIn the indirect design approach, the spudcan penetration resistance is assessed by inserting the extracted soil parameters into the semi-empirical bearing capacity models discussed previously. Extraction of soil strength profiles for different soil types are discussed below.

Figure 8. normalised data plotted on schneider et al.’s [42] soil classification chart

1000

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10

1

Qtn

layer 1

layer 2

layer 3

layer 4

layer 5

layer 6

layer 7

layer 8

layer 9

lc = 2.05

lc = 2.35

lc = 2.76

lc = 3.6

-6 -4 -2 0 2 4 6

U

Figure 7. identified layers through statistical analysis

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identified layers


table 1. identified layers and soil types for the case presented in Figure 4

using CPtu profiles (this study) Field (borehole log) note

layer depth (m) soil layer depth (m) soil

1 0-2 Silty Clay 1 0-2.9 Silty Clay Very soft to soft greenish grey

2 2-2.9 Silt

3 2.9-3.9 Silty Clay 2 2.9-6.9 Silty Clay Stiff greenish grey

4 3.9-5.5 Clayey Silt

5 5.5-6.9 Clay

6 6.9-10.2 Clayey Silt 3 6.9-11.7 Silty Clay Stiff greenish grey

7 10.2-11.3 Sandy Silt

8 11.3-13.8 Silty Sand 4 11.7-15.5 Silty Sand Medium dense light brownish grey

9 13.8-16 Silty Clay 5 15.5-16 Silty Clay Stiff dark greenish grey

Clay: The undrained shear strength (su) is not a fundamental soil parameter since its value for a given stress history depends on the mode of shearing (due to strength anisotropy) and strain rate. From an international collaborative project involving a high-quality database of lightly overconsolidated clays, Low et al. [29] recommended an Nkt,suave value of 13.6 for estimating the intact undrained shear strength from the penetration resistance qcnet; this value was based on either the average su from triaxial compression, simple shear and triaxial extension tests or, in the absence of all three strength data, on the value measured in simple shear. For more heavily overconsolidated clays (increasing Qtn, decreasing Bq) an increased Nkt value of 17.4 can be adopted (Jaksa [47]). Alternatively, su can be deduced using the critical state expression linking OCR to yield stress and vertical effective overburden pressure. After calculating an initial su profile, a simplified linear soil strength profile is then derived for each identified layer, using simple statistics (Lacasse et al. [48]; InSafeJIP [10]).

Sands: For sands (low U), the customary parameter extracted is the relative density, primarily as a function of Qtn. This may be used in an iterative approach to

predict φ, based on an assumed critical state friction angle, φcv (Bolton [49]). A mobilisation function will allow for the finite spudcan displacement required to mobilise the limiting resistance; for highly compressible soil types, such as calcareous sands, the mobilisation function will allow the bearing resistance to be reduced below that corresponding to a minimum (critical state) friction angle. A recent trend is to interpret friction and dilation angle in terms of a state parameter, ψ, instead of relative density (Jefferies & Been [45]). The state parameter combines the influence of void ratio and stress level with reference stress level with reference to an ultimate (steady or critical) state and can be used to describe sand behaviour. Research has shown that the state parameter is a more meaningful parameter to represent the in-situ state sandy soil than relative density.

Intermediate soils: The most difficult soil conditions are intermediate soils (silts), where some degree of consolidation occurs during the penetrometer test, but non-zero excess pore pressures are measured. In most soils of this type, spudcan penetration will occur under undrained conditions, because of the large diameter. Hence it is necessary to correct


in Liyanapathirana [54] and Hossain & Randolph [55]. For trapping of softer material beneath the advancing spudcan, comparing with the results on uniform clay, Hossain & Randolph [56] calculated reductions in the (ideal) bearing capacity factor as ΔNc = 12-11.3 = 0.7 (smooth-based) and 13.1-11.3 = 1.8 (rough-based). These were for spudcan geometries similar to those for Marathon LeTourneau Design rigs and KFELS B Class spudcans. Erbrich [28] estimated the reduction of bearing capacity factor (ΔNc ) due to trapping of softer material beneath a relatively flat-based spudcan (Ensco 102) as around 0.4 in calcareous clay and 0.1 in calcareous silt. Bearing capacity factors for

(b) Effect of soil type and properties on backbone curve

Figure 9. effect of normalised velocity on penetration resistance

the net cone resistance to an equivalent undrained value, according to the backbone curves shown in Figure 9, and based on the normalised velocity (Lee & Randolph [50]):

(8)

where v is the penetration velocity (for spudcans this is often around 0.4~4m/hr), cv the coefficient of consolidation of the soil and De the equivalent diameter of the spudcan. The latter term is the diameter of a hypothetical circle that has the same planar area (A) as the penetrating object. An estimate of the consolidation coefficient, cv, may be obtained from combined correlations of Qtn and U (Hossain et al. [46]). For silts and clay, due allowance must be made for differences in strain rate (proportional to v/D, which is 7.4 × 10-6~1 × 10-4 s-1 for a spudcan and ~0.5 s-1 for a cone), as noted by Erbrich [28].

Direct Cone to Spudcan Design ApproachFor developing direct correlations between cone and spudcan penetration resistance, the first step is to evaluate the spudcan resistance as if the layer in question was infinitely deep, with due allowance for different drainage conditions and strain rates (Erbrich [28]; Lee & Randolph [49]). Further adjustment is then undertaken to allow for the spudcan embedment and layer thickness relative to the object (spudcan or cone) diameter, and also the layer sequence which affects the trapping of softer or stronger material carried down with the spudcan.

Relatively Single Layer Soil or Soil Profile with Minor Strength Variation Both objects in undrained zone: In the undrained regime, the rate of penetration (and hence strain rate) has an important influence on the penetration resistance. Most soils exhibit some viscous effect, which leads to higher undrained strengths for higher applied strain rates. The rate of increase in shear strength (or penetration resistance) per log cycle increase in strain rate can be taken as ~0.12 for clay and 0.03 for silt (Erbrich [28]; Chung et al. [51]; Low et al. [51]; Lehane et al. [53]). Deep bearing capacity factors of cone and spudcan accounting for combined effect of strain softening and strain rate can also be found

V =

vDe

cv

Qv /

Qv, u

nd

rain

ed

V = vde/cv

0.001 0.01 0.1 1 10 100 1000

20

18

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2

0

randolph & hope [26] Kaolin clay

Yiet al. [59] g/p' =140, 17.5

Cassidy [58] silt

(a) Change in spudcan reaction with drainage conditions: partially drained instead of fully drained in sands(after Randolph & Gourvenec [12])

0.001 0.1 10 1000 100000

V = vd/cv

Qv /

Qv, u

nd

rain

ed

Viscous effect ~12% / log cycle increase of

strain rate

vs = 1 m/hr ds = 18 m30

1


spudcan and cone increase with depth from a value at the surface to a deep limiting factor. For a spudcan, a depth of ~1Ds (11~20 m) is required to attain a limiting factor where as for a cone a depth of ~5Dc, i.e. only ~0.2 m, is sufficient. For shallow penetration (critical for heavily overconsolidated clay with a high shear strength close to the mudline), it is therefore necessary to account for appropriate bearing capacity factors (limiting value for cone but reduced value for spudcan) as a function of the relative embedment.

An example on the use of piezocone results for spudcan penetration prediction in clay is demonstrated here using the data obtained at the site referred to in Figure 1. In Figure 10, the estimated bearing capacity derived directly from the net cone resistance is presented. Averaged net cone resistance values within 0.1 × spudcan diameter (equivalent diameter of Ds = 14 m in this example) above and below the depth in question were used in the calculation. Ratio of Nc (for spudcan) and Nkt (for cone), accounting for the effects of previously discussed factors, is used to correlate the cone and spudcan bearing resistance for the full penetration depths. Soil backflow accounting for the advancing spudcan volume was considered. From Figure 10, the spudcan penetration profile predicted directly from the cone resistance is consistent with the measured load-penetration response.

Cone and spudcan in different drainage zone: While correlations between penetrometer resistance and soil properties in sand or clay have been established (Lunne et al. [34]; Randolph [57]), their relationship for intermediate soils and their use for bearing capacity prediction are still very limited. In sandy and silty soils, cone penetration may occur under drained or partially drained conditions while partially drained to undrained conditions may occur for spudcan penetration (Equation 8). A reduction factor should be applied depending on the soil type and corresponding backbone curve (see Figure 9). The current framework is currently being extended to cover intermediate soils, using ideas such as those presented by Lee & Randolph [50].

Layered Soil or Soil Profile with Strong Strength Variation For layered profiles, two other factors must be considered: (a) the differential relative thickness of each layer (t/Dc vs t/Ds; where t is the thickness of a layer), and (b) the trapped soil plug beneath the advancing spudcan base, particularly following penetration through a hard layer into softer material. It is of interest to compare the T-bar and spudcan penetration resistance curves from a centrifuge test and cone and spudcan penetration curve from field data, as shown on Figure 11a. Marked differences are obvious between the two penetration profiles in terms of the overall form and relevant features, largely associated with the differences in size. The peak in resistance within an interbedded strong (2nd) layer is much greater for the T-bar (with DT-bar = 1 m). A simple correlation relating spudcan and T-bar bearing capacity factors, Nc/NT bar, even taking account any differences in strain rate, cannot be used in stratified deposits where strong changes in strength occur. As discussed by Erbrich [28], the large diameter of the spudcan relative to the layer thicknesses allows different penetration mechanisms and a reduced maximum resistance in the stronger layer. This needs to be taken into account when developing penetrometer-spudcan relationships in multi-layer soils with distinctive changes in shear strength.

The thickness of the stiff layer relative to the object diameter is much higher (8 times in this particular model test) for T-bar penetration. The thickness ratio is a crucial factor for assessing punch-through

Figure 10. spudcan penetration prediction directly from qcnet in relatively single layer soil

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stillwater reaction

Preload reaction


(Hossain & Randolph [16, 17]) and the limiting squeezing depth is also a function of the size of the penetrating object (and also the thickness of trapped soil plug carried down in the case of a spudcan; see Figure 11b). As such, the T-bar penetration resistance shows relatively sharp changes at layer interfaces, with a rapid increase near the surface of a stiff layer and a drop at the surface of a soft layer. In interbedded stiff layers, the T-bar resistance rises to attain a local maximum value after a short transition. It then starts to drop rapidly just above a lower soft layer, and continues to decrease for a certain distance into the lower layer. By contrast, the much larger spudcan (Ds = 8 m) tends to integrate the behaviour of the different soil layers for some distance above and below the spudcan, with significant differences in the failure mechanism as it punches through the stronger layer.

After a punch-through the T-bar resistance in the underlying layer drops much lower than the spudcan resistance does. This reflects the fact that (a) a soil plug of significant depth is carried down with the spudcan (Hossain & Randolph [16, 17]; see also Figure 11b), augmenting the penetration

Figure 11. spudcan penetration prediction directly from qt in multilayered soil

resistance through additional resistance around the plug periphery and (b) the soil flow mechanism around the spudcan is impeded by the stronger layer above the spudcan, again increasing the resistance.

To identify the behaviour of spudcan penetration in a wide range of soil types and layering conditions, an experimental programme is being conducted at the Centre for Offshore Foundation Systems (COFS) at the University of Western Australia under a COFS-Keppel collaborative research project. The study covers centrifuge modelling of spudcan installation in intermediate soils such as calcareous sand and silt, as well as silty siliceous sand. The behaviour of spudcans and cone penetrometers in multi-layered soils, including profiles containing intermediate soil layers, is also being investigated. The study is expected to extend the use of penetrometer results to derive more representative shear strength parameters for use with conventional bearing capacity methods and an analytical framework to enable direct correlation with spudcan penetration behaviour for any soil type and layered conditions. Progressive calibration of the method against observations made during rig moves is also being undertaken.

(a) Comparison of bearing pressure from spudcan and T-bar penetration tests

0

20

40

60

80

100

no

rmal

ised

pen

etra

tio

n, d

/d

Vertical bearing pressure, qu (kPa)

0 100 200 300 400 500 600

0

0.5

1

1.5

2

2.5

layer interface

supdcan

ds = 8 m

dse = 8 m

t-bar

dt = 1 m

dte = 2.26 m

soft

stiff

soft

stiffVs = 0.28 mm/s

Vt = 1 mm/

(b) Spudcan with a trapped plug after a punch-through failure (axes in mm, model scale)

-60 -40 -20 0


Figure 12. schematic diagram of integrated jack-up installation monitoring system (Quah et al. [60])

Field Penetrometer data

geotechnical interpretation Direct penetrometer-spudcan

Shear strength derivation

Design Shear Strength from other tests

Target Preload Base Reaction from SSA

identification of Punch through Potential

Punch-through?

Within tolerance?

expected Penetration bounds

Preloading Plan

Preloading Monitoring

Prediction of spudcan load-Penetration response

real-time data Acquisition • Total water depth • Chord load • Draft / airgap • Leg reaction • Leg marking

evaluation of structural implications/P-t simulations

Preloading strategy Safe draft/airgap, sequential legs

Yes

No

No

Yes

Penetrometer resistance, shear strength

Predicted penetration, punch-through behavior

Predicted and observed load-penetration response


AN INTEGRATED SYSTEM FOR JACKUP INSTALLATION The discussed design approaches are to be incorporated into an integrated system for jackup installation. The schematic diagram of the system is illustrated in Figure 12. The system is aimed primarily at identifying punch-through potential and closely monitoring progression of spudcan installation relative to predictions. The system consists of four sub-systems: spudcan penetration prediction system (SPPS); punch-through simulation system (PTSS); preloading planning and stability system (PPSS); and leg penetration monitoring system (LPMS). The sub-systems are described briefly below.

Sub-System 1: Spudcan Penetration Prediction SystemThe sub-system automatically calculates the spudcan load-penetration curve from penetrometer results or a set of soil shear strength parameters. There are two approaches that can be adopted to obtain the spudcan penetration curves: i) direct correlation from penetrometer result; ii) calculation based on predefined soil layer boundaries and soil parameters using semi-empirical bearing capacity models. The soil parameters in the latter approach can be derived directly from the penetrometer results or other shear strength tests. Along with the target preload base reaction derived from in-place analysis conducted at the site-specific assessment stage, potential punch-through can be identified from the predicted spudcan load-penetration curve.

Sub-System 2: Punch-Through Simulation SystemThe sub-system is essential to estimate the behavior of the leg when experiencing punch-through under given initial conditions. Load-penetration response of the leg as well as structural integrity of the leg and holding system during a punch-through event can be simulated with PTSS. The maximum punch-through depth, at which the leg reaction is again equalized by the soil bearing capacity and buoyancy of the hull, can be estimated based on the predicted load-penetration curve. If there is any structural over-utilization indicated from the analysis, alternative preloading strategies such as preloading ‘underwater’ (partially submerged with draft condition) or preloading with zero airgap, can be implemented to ensure that the installation is carried out within the

rig tolerable limits. Based on the understanding of punch-through behavior, the jackup operator can assess, anticipate and mitigate the risk of potential punch-through at the installation site.

Sub-System 3: Preload Planning and Stability SystemThe sub-system is developed to assist the jackup operators to plan for preloading operations in order to achieve the target preload base reaction. The system allows planning of various operation scenarios to meet preload requirements, including simultaneous preloading or sequential preloading, staged preloading, preloading with or without air gap, or preloading “underwater”. PPSS is developed in such a way that the jackup operator is able to input and retrieve the preload tanks level conditions at any given preloading stages.

Other than planning of the preloading operation, PPSS can also serve as a stability interface during rig transit conditions. With this stability interface, jackup operators can configure the required ballast water in each ballast tank in order to keep the hull stable during rig transit.

Sub-System 4: Leg Penetration Monitoring SystemThe sub-system monitors both leg penetration and reactions during jacking and preloading operations on a real-time basis. This enables the jack-up operator to evaluate the observed leg-penetration response against the predicted leg penetration curves and to be alerted when the legs are approaching a critical penetration depth at which rapid leg penetration is predicted to occur. Should any unpredicted behavior observed, necessary measures can be taken in time as installation is being monitored continuously.

In the monitoring of leg penetrations, LPMS takes into account several input parameters such as leg flag marking readings, distance from the hull base to the leg flag marking, water depth (including tidal variation), and also air gap or draft level at each of the legs. It should be noted that accurate reading of leg penetration will be hard to achieve without proper sensors for measuring water depth, draft, and air gap. The necessary advanced measurement technology has been extensively used in the marine industry for some while, despite limited use for jackups.


To monitor the leg reactions, LPMS accounts for distribution of total weights on the jackup, including the weight of the hull, cantilever, drill floor, solid variable loads, liquid variable loads, preload ballast, legs and spudcans, together with their respective lateral, transverse and vertical centres of gravity. Buoyancy effects, due to submerged part of the leg/spudcan and partial hull submersion under draft conditions, are also incorporated in the system. With proper account of the buoyancy force, both leg load and leg reaction can be monitored accurately.

CONCLUSIONDepletion of reserves in traditional and shallow waters regions is resulting in exploration in deeper, unexplored and undeveloped environments with more complex seabed soil conditions. Installation of jack-up rigs in complex stratified seabed, where an interbedded strong layer overlays a weak layer, is challenging due to the potential for punch-through

failure. An accurate spudcan load-penetration curve is therefore essential to ensure safe jackup installation.

This paper has described design approaches to obtain spudcan load-penetration curves based on penetrometer results. The design approaches include an indirect method, by incorporating the extracted soil parameters into semi-empirical bearing capacity models, and a direct method relating the spudcan resistance more directly to the cone resistance. The design approaches are to be implemented into an integrated jackup installation system that comprises separate modules for spudcan penetration prediction, potential punch-through identification, preload planning, and preload monitoring. Combined with existing data acquisition systems typically available on-board jackups, such an integrated system is capable of providing the necessary information at every stage of jackup installation, assisting jackup operators in the safe installation of jackups.


ACKNOWLEDGEMENTSThe authors would like to thank Steve Nowak of Integrity Offshore for provision of the field data. Dr Chen Hong’s work on the development of leg punch-through simulation software is greatly appreciated. The research presented was undertaken with support from the Australian Research Council and the industry partner Keppel Offshore and Marine, Singapore through the Linkage Project (LP110100174). The work forms part of the activities of the Centre for Offshore Foundation Systems (COFS), currently supported by the State Government of Western Australia as a Centre of Excellence, and now forming one of the primary nodes of the Australian Centre of Excellence in Geotechnical Science and Engineering. This support is gratefully acknowledged. The authors would also like to acknowledge the organizer of the 13th International Conference: The Jackup Platform (City university London, uK) for the permission to reproduce this paper.


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[43] Lunne T, Andersen KH (2007). Soft clay shear strength parameters for deepwater geotechnical design, Keynote Address. Proc. 6th Int. Offshore Site Investigation and Geotechnics Conf.: Confronting New Challenges and Sharing Knowledge, London, uK, 151-176.

[44] Mayne PW, Peuchen J, Bouwmeester D (2010). Soil unit weight estimation from CPTs. Proc 2nd Int. Symp. Cone Penetration Testing, CA, uSA.

[45] Jefferies M, Been K (2006). Soil liquefaction, a critical state approach. Taylor and Francis, London and New York.

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[47] Jaksa MB (1998). An experimental study to quantify the cone factor of a stiff, overconsolidated clay. Australian Civil Engineering Transactions CE 40, 43-48.

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[49] Bolton MD (1986). The strength and dilatancy of sands. Géotechnique 36, No. 1, 65-78.

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[52] Low HE, Randolph MF, DeJong JT, Yafrate NJ (2008). Variable rate full-flow penetration tests in intact and remoulded soil. Proc. 3rd Int. Conf. on Geotechnical and Geophysical Site Characterization, Taylor & Francis Group, Taipei, Taiwan, pp. 1087-1092.

[53] Lehane BM, O’Loughlin CD, Gaudin C, Randolph MF (2009). Rate effects on penetrometer resistance in kaolin. Géotechnique 59, No. 1, 41-52.

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Installing Offshore Wind Turbines in Harsh Environments 113


this papeR outlines how wind tuRBine Blades may Be handled on a laRge fouR-legged JacKup Vessel, and installed safely and efficiently on the offshore location using the Principle of Horizontal Guiding for all lifting operations. A Knuckle Boom Crane, with blade gripper in its wrist, is a central component for the layout of topside equipment for this vessel. Specifically, it is equipped with a racking system to store blades as triples in a common frame, which fits the blade cradles commonly in use for supporting the blades from factory to installation site. The cradles fit inside this frame, so safe and consistent handling of blades can be obtained, independent of chosen blade supplier, in the harbour and on the vessel, and when laying down empty cradles inside the frames on the vessel and when returned to the harbour.

Blades are assembled by the Knuckle Boom Crane to rotors in a low elevation on a fixture turning the rotor’s hub, before the complete rotor is hoisted to the top of the Wind Turbine tower. Two alternative means of handling the rotor is shown, all fitting the same layout with blade rack and Knuckle Boom Crane. The vessel layout also involves a lattice boom crane, for handling of tower sections and nacelles from the harbour to the vessel.

Installing Offshore Wind Turbines in Harsh Environments

asbjorn moRtensen, Dr. Ing, M.Eng Keppel Offshore & Marine Technology Centre

sudhan Vincent, B.Eng Keppel Offshore & Marine Technology Centre

presented at ewea offshore conference 2011


rotor blade (3)

nacelle

hub

tower sections (1 or 2)

Foundation

INTRODUCTION The offshore wind industry is based on the same Wind Turbine (WT) as widely used in onshore applications. The most common assembly method in use follows a sequential procedure, starting with one or two vertical Tower sections bolted to a grounded Foundation. Then follows the Nacelle housing the generator and gearbox, located on top of the tower. Finally follows the Rotor, which consists of a Hub and three Blades.

installation, with capacity to carry 5-10 WT’s for each harbour visit. The described assembly method in this paper is based on the conservative fact that most WT’s are assembled from individual parts having lifting points certified for crane lift, and with an existing lifting technology using spreader beams, yokes, grippers, fixtures/cradles, and special assembly tools as hub tilting/turning tool.

Rotors are typically assembled in two ways:

1. Elevated assembly, horizontal blade/hub turning method: The nacelle has a hub turning motor, enabling the hub to be slowly turned 120 degrees before the next blade is bolted into the hub. A crane with a blade gripper lifts the blade from the ground to the top of the WT, always in horizontal orientation. The blade gripper may have integrated blade rotation and blade pitching features, and typically using two wide canvas belts to strap the blade from above.

2. Ground assembly, horizontal rotor method: The 3 blades are bolted to the hub sitting on a turning fixture on the ground, then one or two cranes lift the assembled rotor to the top of the WT, using a hub tilting lifting device, and bolts it into the nacelle shaft.

This paper describes a method to store a multitude of blades on a JU vessel, suitable for both of these assembly methods. The rotor is assembled on the offshore location, in a lower elevation than the top of the WT tower, by using either of these two methods.

Although a number of other assembly methods exists, like the “bunny ear” method, harbour assembly and erect/slanted transportation for offshore installation, the two methods are proven. At the same time, it is a growing perception in the industry, that safe lifting operations is essential, to improve the accident track record of the industry. As the industry grows with increasing WT sizes, it

A number of very innovative assembly methods for offshore WT have been proposed in the literature, in order to make its assembly more cost efficient and having less downtime due to weather conditions. The methods are strongly dependent on the chosen vessel configuration. This paper describes the use of a large Dynamic Positioned (DP) four-legged Jackup (JU), in order to create a stable platform for

When developing improved assembly methods for harsher weather, the key challenge is the need to slightly change or re-certify lifting points on the various WT parts.

Figure 1. Wind turbine parts


is important to develop generally accepted methods. When developing improved assembly methods for harsher weather, the key challenge is the need to slightly change or re-certify lifting points on the various WT parts.

BLADE RACKING AND WT ROTOR ASSEMBLYAn assembled WT rotor occupies a large volume when stored on a vessel. Furthermore, due to its large diameter up to 150 m for 7-10 MW turbines, it is difficult to transport such rotors out of most harbours on a vessel. Special transportation solutions may be developed, but such vessels often lack the flexibility to be used for other offshore operations, and are therefore a commercial risk for the owners. When blades are stored aside each other, aligned in parallel, they occupy far less space. This paper shows an elevated racking solution with blades stored athwart ship. The functional idea is to design a racking and crane solution with minimal distance between blade storage position and assembled rotor position, and to guide all hanging, swinging loads so high uptime for wind can be achieved. The application of this idea must include easy human access to all points of operation and maintenance, hence a high level of safety standard and efficiency is achieved.

Crane and gripper used for blade assembly to a rotorBlades are relatively lightweight compared to nacelle and tower sections. They have a high wind catching area/ weight ratio, due to their intended use, meaning even low wind speeds will cause large side forces causing the blade to swing if freely suspended. A governing principle to enable safe blade handling, is to use a Knuckle Boom Crane (KBC) with a wrist mounted gripper. This is a proven crane solution used in offshore drilling operations, which is certified for handling tubulars in severe windy weather, like a storm. The key is to avoid using a standard lattice boom crane with hook suspended from a hoist wire, as this creates swinging motion

The application of this idea must include easy human access to all points of operation and maintenance, hence a high level of safety standard and efficiency is achieved.

Figure 2. Knuckle boom Crane with wrist mounted yoke-based gripper

of the blade. Use of horizontal tugger winches to control those motions create dangerous operations, or time consuming operations.

The yoke is bolted to the native gripper supplied from the blade supplier, if available, or the yoke carries two canvas belts with remote quick release system (when unloaded).

This KBC wrist comes with a hydraulic/electric swivel for power and signals, and a powered slew motion so the blade rotation can be controlled in the horizontal plane (to counteract wrist torque created by the wind). The wrist mechanism is like a universal joint with one or two brakes for the two additional tilt motions, and is preloaded by gravity so the blade tilt is always hanging straight down when the brake is off. Since the pivoting

Knuckle boom Crane with operator Cabin

Wrist with slewing Yoke

empty blade rack w/ ladders

blade gripper w/ Canvas belts


point is close to the blade, and gravity is preloading against wind forces, this freedom of motion is not severely affecting the blade motion in strong wind, besides, some freedom of motion is necessary when picking and installing the blade. The blade gripper typically comes with a canvas belt tension control so the blade pitch can be regulated prior to inserting the blade into its hub location. The yoke can also be equipped with blade longitudinal motion, to ease the blade insert operation, or to intentionally tilt up or down the blade when matching its hub flange with the hub orientation. This type of gripper and its motion control follows the same design principles like tubular grippers used for oil drilling rigs.

Depending on the nacelle and hub technology provided by the WT provider, hub pitch and hub rotation may be remotely controlled from the nacelle, as a trade off instead of providing blade pitch and blade longitudinal motion+tilt to achieve the same from the KBC, when matching the blade bolt pattern and the hub end flange.

Blade handling in the harbourIn order to reduce loading time for the vessel in the harbour, blades should be handled three by three in

a common fixture from quayside to the vessel racking system, see Fig. 3. Since blades are stacked high on large vessels, it is important to have a safe stacking principle to avoid dropped objects and other dangerous handling situations. Each blade comes from the factory in a root end cradle and a tip end cradle, stackable similar as ISO containers, or similar system. These cradles are lifted by use of a spreader bar and two frames suiting triple blades, in the harbour. Since blades are matched in the factory, the blades belonging to the same WT must be handled together. If any blade shows or develops a failure during the offshore assembly, all blades in the set must be set aside and brought onshore for warranty work. This is an important functional requirement for a blade racking system.

Blade Cradle handling in FramesThe cradles supplied from the blade manufacturer, should fit into triple frames with walkways, which can be stacked on the vessel. It is an advantage to split such access equipment functionally apart from the cradles, to reduce the amount of harbour work when preparing the cradles for next operation (either go offshore, or return to factory). For general information, such cradles may be 2-3 m high and wide, or more, and heavy, due to the size of the

Figure 3. triple blade handling frames

spreader bar

Frame with root end Cradles

Frame with tip end Cradles

blades. There have been accidents when handling empty cradles; for example, to lay down the cradle after lifting off the blade, may be tricky if not prepared properly.

Blade storage in racks on the vesselThe triple blade frames fit into slots in a blade racking system on the vessel. Note that the rack has ladders and walkways available to all areas of human access, i.e. those points where the blades are bolted or suspended into the cradles.

Note also the elevated position of the blades, which is necessary due to the next step in the installation process, when assembling the rotor from 3 single blades. The rotor can be assembled to the hub in vertical or horizontal orientation. In either case, the hub must be elevated so the rotor can index for each new blade to be added from the same position by use of the KBC. See section on about rotor indexing.

Optional: Nacelle storage under the blade rackNacelles can be loaded into this deck space from above, when in the harbour, prior to the blades are stored in the rack above. During offshore assembly of turbines, one by one nacelle is pulled out by use of hydraulic driven tractors. The tracks are elevated from the main deck, so these profiles can be moved, and width can be adjusted when rearranging for another vendor’s nacelle.

Figure 5. blade racking system, elevated

Figure 4. root end cradles from supplier, slotted into blade frame from

root end Cradles empty (3)

blade Frame w/ walkway



Cherry picker with manriding basket for blade gripper accessIn order to reach the two canvas belts hanging down from the blade gripper, prior to wrapping them around the blade, it is necessary to provide human access to pick up the canvas and pull it up into the tensioning system. See the following picture showing two manriding baskets mounted on telescopic cherry pickers. The blades are always picked in the same sequence, so the baskets move above the next level of blades. Other access systems are also possible.

The canvas belts shown have a certified automatic release system, working only when the belt tension is nil, to avoid accidentally dropping the blades. The gripper is released after the blade has been safely bolted to the hub.

Rotor assembly in low elevationThe rotor assembly in a low elevation requires a stable platform in order to enable the hub rotation, by use of a dedicated fixture with a rotation device, or by use of hub turning device from the WT supplier. Various solutions to support the hub are shown.

It is necessary to give access for human operators inside the hub, in order to attach nuts to threaded bolts protruding from the blade flange (common design detail for most blade suppliers).

WT ROTOR INSTALLATION, LAST STEP IN SEQUENCEThis paper has described how rotors can be safely assembled in a reduced elevation on an offshore jackup in high wind speeds, typically up to 12-15 m/sec or more, depending on local strength of WT parts and their lifting points. Cranes may be certified up to 20 m/sec. The vessel topside equipment involves fixtures or a mast structure attaching the hub to a stable point, to include a device for local rotation, such that the rotor can swing around without hitting other objects like or the jackup legs, and such that human operators can safely enter inside the nacelle or hub. The final movement of the rotor to the top of the WT can be enabled by a crane, or by using a mast with winches for guided hoisting, and a cantilever extension system to achieve position control. See below for details.

Figure 6. nacelles stored under the blade rack, on skidbases on an XY skidding track, with tractors, also showing telescopic walkway to the foundation in the front

Figure 7. Manriding baskets to access the canvas gripper when grabbing the blades

Figure 8. Vertical rotor assembly in low elevation, using torque fixture to index the hub



Figure 10. Vertical rotor assembly in low elevation, using KbC and nacelle with hub turning device

Figure 9. horizontal rotor assembly in low elevation, using KbC and hub turning device


distinct elevations, it is shown driven forks that engage with the torque motor which rotates the rotor. These forks extend and provide a stable support for the rotor during single blade installation, and when doing the final turning to align the rotor lifting tool so one blade is pointing vertically down.

The mast shown in Fig. 11 sits on a cantilever able to extend out above the sea, close to the foundation. The mast has transverse movement, enabling an accurate positioning of the rotor prior to attachment to the nacelle. The mast has tugger winches at various elevations, used to guide the rotor to avoid swinging motion during the installation. At two

Figure 11. guided installation of rotor by use of cantilevered mast, and hub torque motor temporarily elevated out of way


The vessel owner operating the WT installation vessel, faces many challenges when ordering newbuilds. One major concern is of course the cyclical nature of a wind energy business still in its infancy years, uncertainty created by WT technology innovation, coupled to many years delivery time for jackups. This will drive the owner building a vessel on speculation, to configure the cranes creating high

Figure 12. rotor installation by use of hub-tilting device hanging in a crane hook

When using a cantilevered mast to install the rotor, providing XY position control above the foundation, it is using a dolly running up and down in a track bolted to the vertical mast. This dolly supports against swinging from horizontal forces affecting the part under installation, at the same time as it is lifted from above. This ensures minimum downtime during installation in windy weather.


flexibility, fitting also other offshore applications. For the wind farm owner, not in the business of operating or owning vessels, he wants to lease an installation vessel with minimum downtime due to weather, and have a deck load capacity and space fitting the chosen WT supplier he has contracted for his wind farm. A major efficiency and safety factor for WT assembly, consuming a lot of deck space, is rotor and blade handling. Blades are huge, up to 75 m long or more for 5-10 MW generators, and they are handled from the factory via the road and harbour, to the offshore location, in frames or cradles.

Wind turbine vendors should certify their standard lifting points to withstand such higher wind speeds, and such combined lifts as shown here. The advantage of the described methods are that it is used a standard arrangement of lifting points, and standard sequence for the assembly of the

turbine. The vessel size fits large 5-10 MW turbines, and the hull can carry a high number of wind turbines, suitable for remote installation sites far from the harbour.

CONCLUSIONIt is shown how offshore wind turbines can be installed in high wind speeds, by using the Principle of Horizontal Guiding during all lifting operations. The vessel’s topside equipment is designed for safe use by human operators, to all areas requiring access.

The main enabler of this principle is to use a Knuckle Boom Crane to lift blades, to provide horizontal guiding from the wrist of the crane, so high wind speed does not cause downtime. An affordable crane’s limited reach envelope in turn leads to the use of an elevated rack for storing single blades, so blades can be assembled to rotors in a lower elevation than the top of the wind

Figure 13. guided rotor and nacelle installation by use of cantilevered mast


ACKNOWLEDGEMENTS

The authors would like to acknowledge the European Wind Energy Association (EWEA) for the permission to reproduce this paper, originally presented at the EWEA Offshore 2011 conference, held in Amsterdam, 29 November – 1 December 2011


REFERENCES [1] K.S. Foo, A. Mortensen, T. T. Wong and Y. J. King, “Offshore Wind Turbine Installer”, KOMtech Technology Review (2010). Pg 69 – 74.

turbine. The final enabler for the last step of the installation process, in order to close the loop and always follow the principle of horizontal guiding, is to use a cantilever and mast when bringing the nacelle with the rotor on top of the WT tower.

It is shown three alternative layouts fitting the same blade rack and crane layout, depending on how windy the final wind park destination is, and depending on the rotor tilting and nacelle hub turning technology available from the turbine supplier. For the benign installation weather case, a simple rotor turning device can be used, using a lattice boom crane for the final rotor handling to the top of the tower. For the more windy weather

locations, it is shown how a cantilever with a transverse skidding mast and a track with a guide dolly, can be used to guide the nacelle and rotor, or the hub/rotor alone.

It is recommended to plan future large wind parks using such installation techniques. For the wind park owner, prior to the ordering of new turbines, it should be dedicated vessel and assembly method, so the turbine’s lifting points including blade cradle interfaces, and hub turning technology, can be planned in detail. This is necessary in order to minimize downtime due to weather during the installation period, and accurately predict overall installation cost.

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New Generation Deep Water OSVs for Oil & Gas Operation 125


New Generation Deep Water OSVs for Oil & Gas Operation

tan cheng hui, C.Eng, P.Eng, MSc, BSc Keppel Offshore & Marine Technology Centre au yeong Kin ho, MSc (Mech), MSc (Tech Mgt), B.Eng Keppel Singmarine

presented at osV 2011 conference

offshoRe suppoRt Vessels (osVs) play a Vital Role in maintaining and suppoRting offshoRe oil and gas (o&g) opeRations. In recent years, as shallow water oilfields have been largely exploited, O&G companies are moving rapidly into deep water O&G exploration and production. To support the more complex deepwater field developments, new generation of OSVs, in particular Platform Supply Vessels and Anchor Handling Tug Supply Vessels are required. These vessels are not only larger in size, capacities and power, they also have better capabilities in terms of automation, position keeping, manoeuvring, seakeeping, and cargo handling. This paper gives an insight into the features of new generation PSVs and AHTS.

In addition, the paper describes two proprietary designs developed by Keppel Singmarine’s technology unit Marine Technology Development (MTD). They are MTD9045-DE, a 4500 Dwt PSV and MTD80210A, a 18000 Hp AHTS capable of bollard pull in excess of 210 tonnes. These designs are tailored to Petrobras’ requirements to support deepwater exploration and production in the pre-salt layer offshore Brazil, which have specific requirements on both anchor handling equipment and supply capabilities.


Anchor Handling Tugs (AHTs) – outfitted with large winches and cranes to set and lift anchors, towing of mobile offshore installations such as semi-subs, jack-ups and barges

Anchor Handling, Towing, Supply Vessels (AHTS) – AHTs with large deck space to deploy a range of equipment required for oilfield exploration

Field Support Vessels (FSVs) – designed to support tasks such as high speed supply runs, crew transfers, safety standby, inshore and offshore survey activities

Multi-Purpose Support Vessel (MPSVs) – designed to provide a variety of subsea support services, including ROV support, subsea construction, post-drilling well services

Specialised vessels – Vessels operating in niche markets such as cable-laying, pipe-laying, rock-dumping, diving support, geophysical, well-stimulation, well-intervention, accommodation vessels

There are about 6000 OSVs worldwide,of which two-third are AHTS and PSVs.

NEW GENERATION PSVs AND AHTSTo meet the rising demand for more complex deepwater field developments, new generation PSVs and AHTS are becoming larger in size, capacities and power. They have evolved into highly sophisticated support vessels, capable of undertaking a multitude of additional complex tasks, including diving support, subsea construction support, inspection, maintenance & repair, well stimulation, fire-fighting, oil recovery, and stand-by rescue. These additional functions require specialized equipment and operating staff, and sophisticated dynamic positioning systems.

FEATURES OF NEW GENERATION PSVs• Specialisedvessels(eachPSVperformsaspecific duty and carries large amounts of few products)

• Increasedcargocarryingcapacity(greaterthan 4500T deadweight and large mud tanks)

• Largedeckspace(>1000m2)

• Increasedtankcapacities(Hugecapacities for water and fuel oil)

INTRODUCTION Offshore support vessels (OSVs) play a vital role in maintaining and supporting offshore oil and gas operations. In recent years, the oil fields for shallow water drilling have largely been exploited, this gives rise to oil and gas companies moving into deeper waters, especially the Gulf of Mexico and off the coasts of Brazil and West Africa, in their search for oil and gas. There are also new deepwater discoveries in South-East Asia. Although there is no industry consensus for the definition of deepwater oil fields, it is commonly refer to those below more than 300 metres of water. Much of these deepwater oil and gas fields lie in complex geologic formations and they are hidden beneath a mile or more of salt layers.

Deep water oil & gas exploration and production (E&P) require large number of floating drilling and production units. This in turn creates strong demand for OSVs designed for installation, supply, anchor handling and maintenance.

Supplies for offshore E&P include potable water for crew consumption, fuel oil for generators, drilling mud, brine, methanol, dry bulk (cement, barite, bentonite), deck cargo (drill pipes, casings, risers, drill bits, collars, blow-out preventers and offshore containers). As these OSVs carry increasing quantities of oil products, hazardous and noxious liquids, specific safety measures also need to be taken into account in terms of design.

The continuous expansion of the oil & gas related offshore activities in other regions, in particular in the Arabian Gulf, South East Asia, Mediterranean, Black Sea & Caspian Sea, as well as the Arctic & sub-Arctic regions, also pushes the OSV market forward.

THE DIFFERENT OSVsThe OSV market has become very diversified, with more than a dozen different niches within this category. The market now looks at their specific capabilities:

Platform Supply Vessels (PSVs) - designed specifically for carrying supplies to and from offshore installations


• Flexibleconfigurationwithrespecttoliquid and bulk cargoes

• Excellentmanoeuvringcapabilities

• Advanceddynamicpositioning(DP)systems for superior station keeping. Minimum of DP Class 2 redundancy.

• State-of-artautomationandpowermanagement system for onboard electrical systems for better and more accurate controls. Also to facilitate minimum manning.

• Dieselelectricpropulsionforlowerfuel consumption, especially during DP operations.

• Diesel-electricarrangementstoprovidegreater flexibility to the internal arrangement. The smaller footprint of the engine room offers alternatives for optimising the space for supporting rig operations such as increasing liquid mud tank capacity.

• Advancedcargohandlingsystemstoenable faster loading and discharge of dry bulk and liquid cargoes.

• Panoramicnavigationbridgehaving360degree view for enhanced operational capability.

• Rollstabilizingtankstoreduceshipmotionsin harsher sea environment.

• Largeaccommodationspaceswithenhanced crew amenities

• ComfortClassnotation(lownoiseand vibration, temperature and humidity control) to facilitate crew having to work in more demanding and uncomfortable environments

The largest PSVs are also suitable for subsea construction work and are normally arranged with large accommodation, ROV hangar, moonpool and helicopter deck.

FEATURES OF NEW GENERATION AHTS• Highbollardpull.Indeepwater,highbollard pull in excess of 200 tonnes is used for shifting and for movement of semi submersible rigs in oil fields

• Powerfultowingandanchorhandlingwinches

• Abundantstoragecapacityforfibreropeand wire/chain

• Enhancedstabilityandtrimtohandlethe larger and heavier anchors and chains. Norwegian Maritime Directorate (NMD) has issued Guidelines on the implementation of specific measures to ensure a sufficient safety level during anchor handling operations

• Safeanchorhandlingsystemforenhanced safety for working on deck

• Torpedoanchorhandlingsystems

CLEANER SHIPSThe new generation of vessels have the following features to meet the latest environment demands:

• Double-bottomanddouble-sidedhulltominimize environmental impact of hull penetrations (MARPOL Annex I Oil Fuel Tank Protection and carriage of Noxious Liquid Substances)

• EnginesrunningonMGOtominimize SOx emissions

• Airconditioningplantrunningonrefrigerants with zero ozone depleting potential

• Incineratorplantforburninggarbageatsea

• Sewagetreatmentplantfortreatingblack and grey water prior to overboard discharge

• Tankwashingsystemformudtanks

• Ballastwatermanagementsystem

• Greenpassport

To meet the rising demand for more complex deepwater field developments, new generation PSVs and AHTS are becoming larger in size, capacities and power.


MTD80210A, a 18000Hp AHTS capable of bollard pull in excess of 210 tonnes. The PSV and AHTS designs are specifically developed for the Brazilian market and hence significantly different from KOMtech’s 3,500 dwt PSV design and 160tbp AHTS, which was intended primarily for the general market.

MTD 9045-DE – 4500 Dwt Platform Supply VesselThe MTD 9045-DE is a large-sized PSV, measuring 94.2 metres in length and 19.8 metres in width, having a deadweight capacity in excess of 4500 tonnes and design speed of 15 knots at 4.5 metres draft. The vessel’s internal tanks arrangement and systems are uniquely designed to be capable of functioning as a dedicated fluid carrier (Fluideer) or dedicated diesel oil carrier (Oileer), in both cases meeting the specific requirements of Petrobras. As a Fluideer, the tanks are capable of carrying oil-based mud (1120 m³), water-based mud (520 m³), N-Paraffin brine (260 m³), drilling brine (1490 m³) and dry bulk (330 m³). As an Oileer, the vessel is capable of carrying diesel oil (3500 m³) with deck cargo capacity of 2200 tonnes. The characteristics of this unique cargo is highly valuable for the ship owner as it offers a high degree of flexibility in meeting chartering requirements.

MTD9045-DE is able to accommodate 30 crew members and is equipped with dynamic positioning Class 2 capability. The vessel will be optimally equipped for wide range operation conditions with a diesel-electric propulsion system, consisting of four main generators each rated at 1580 ekW, combining them with two azimuthing drives of 2,500 kW and two 1000 kW tunnel thrusters. The vessels bears the Class notation : ABS, +A1(E), Offshore Support Vessel, +AMS, +ACCU, +DPS-2, UWILD.

Keppel Singmarine’s technology unit, Marine Technology Development, has responded to Petrobras’ requirement with two proprietary designs: MTD9045-DE, a 4500 Dwt PSV and MTD80210A, a 18000Hp AHTS capable of bollard pull in excess of 210 tonnes.

BRAZILIAN REQUIREMENTS FOR OSVsOver the past few years, Brazil has made several huge crude oil reservoir discoveries, which propelled her to become a major player in the world energy market. The first of Brazil's massive pre-salt oil fields, Tupi, was discovered in 2007. This field is located 250km offshore contains between 5 billion and eight billion barrels of oil, is the biggest discovery in the last 20 years worldwide. The Carioca field, located 280km offshore was discovered a year later, and is potentially four times larger and the third largest discovery in history.

Brazil’s huge quantities of high-quality oil are trapped in deep water around 4000 metres under the sea floor, beneath layers of salt and rock. There is also around 2,000 metres of seawater between the seabed and the surface, further complicating the oil retrieval process. This leads to specific requirements related to both anchor handling equipment and supply capabilities.

Brazil is now the world’s biggest market for goods and services in the offshore oil industry, with Petrobras as the biggest single buyer. Petrobras has firm plans to call for tenders of up to 146 OSVs by 2020 to facilitate the E&P of the pre-salt field in the Santos Basin. These ships have to be Brazilian flagged and be built in Brazil as well as to be on long-term contracts. This bonanza includes 64 AHTS, 64 PSVs and 18 oil recovery vessels, are expected to generate a shipbuilding boom in Brazil. Of these, more than 40 large-sized PSVs are in demand by 2014.

Proprietary OSV Designs for Brazilian Market Keppel Singmarine’s technology unit, Marine Technology Development has responded to the Petrobras’ requirement with two proprietary designs: MTD9045-DE, a 4500 Dwt PSV and


Figure 1. Mtd 9045-de – 4500 dwt Platform supply Vessel

Figure 2. Mtd 80210A – 210t bollard Pull Anchor handling / supply Vessel



MTD80210A is able to accommodate 24 crew members and is equipped with dynamic positioning Class 2 capability. The vessel will be optimally equipped with a diesel-mechanical propulsion system, consisting of four main engines rated at 2 x 3840 bkW and 2 x 2880 bkW, combining them with two controllable pitch propellers in fixed high thrust nozzles. Excellent manoeuvring and stationkeeping capabilities are achieved with 1 tunnel thruster and 1 azimuthing thrust each rated for 900 kW at the bow, and 1 tunnel thruster and 1 azimuthing thrust each rated for 900 kW at the stern. The vessels bears the Class notation DNV +A1(E), Tug, Supply Vessel, EO, SF, Clean, Dynpos-Autr (DP2), DK(+), HL(2.0) or equivalent ABS notation.

PSVs in VogueTo date, the MTD 9045-DE design has been adopted for two vessels for Guanabara Navegacao Ltda. The Guanabara PSVs are pioneer batch of newbuilds for Keppel’s new ship-owning division in Brazil. Unlike AHTS, the lack of rampant speculative building during the 2006-08 boom has sheltered the PSVs segment from oversupply.

Both the MTD 9045-DE and MTD 80210A designs have also been selected by a few international offshore shipping companies for Petrobras tender.

MTD 80210A – 210T Bollard Pull Anchor Handling / Supply VesselThe MTD 80210A is a large-sized AHTS, measuring 88 metres in length and 22 metres in width, having a deadweight capacity of 3500 metric tonnes and design speed of 15 knots at 6 metres draft. The tanks are capable of carrying drilling brine (1400 m³) and marine gas oil (1400 m³) with deck cargo capacity of 1700T.

The distinct features of this vessel are the large array of deck machineries which are way above that found on most large AHTS. These machineries are uniquely arranged for operational efficiencies and to meet the very stringent deepwater anchor handling and towing requirements of Petrobras. These consist of double drums main towing/anchor handling winch rated at 400 tonnes line pull, single drum special towing winch rated at 400 tonnes line pull, double drums secondary towing winch rated at 130 tonnes line pull, double drums storage winch, four large storage reels for 208mm dia rope, various sizes of wildcats for anchor chain handling, 650 tonnes SWL shark jaw, tow pins, split-type stern roller, pennant winder and torpedo handling device. In addition, a ROV hangar and ROV launch & recovery system.

Emulsified Fuel System for Marine Diesel Engines 131


Emulsified Fuel System for Marine Diesel Engines

Jerry ng Kok loon, PhD, B.Sc Blue Ocean Solutions (Blue Ocean Solutions Pte Ltd is a 70% owned subsidiary of the Keppel group)

this papeR pResents emulsified fuel system as a compelling solution to ship owneRs foR Reducing fuel consumption. It explains the secondary atomisation effect of emulsified, fuel which creates a better fuel-air mixture for more efficient combustion. The differentiating features of Blue Ocean Solutions Emulsified Fuel System (BOS EFS) and its key innovation in emulsified fuel technology were described. One important practical feature is the ease of retrofitting BOS EFS to existing ships without modifying the diesel engine. The ship does not need to dry dock to be retrofitted; there is near zero operational down time. The experiences and references of BOS span over 20 years, covering RoRo ferries, bulk carrier, container ship and cruise ship powered by slow and medium speed engines. Wartsila has also tested BOS EFS under project Hercules. All installed systems have consistently delivered more than 3% reductions in fuel consumption with no adverse effects to the maintenance of the hot components of the engine.


INTRODUCTION While the cost of marine fuel oil has increased by more than 33% since 2011 (Fig 1), freight rates continued to be depressed by over capacity and sluggish economic growth. Today, bunkers represent about 70 per cent of an average vessel’s total costs, according to the Baltic and International Maritime Council, or Bimco, the Bagsvaerd, Denmark-based organisation that groups two- thirds of the world’s ship owners. For example, Maersk recently reported that its annual fuel cost is US$6 billion. 3% improvement in fuel efficiency translates to a US$180 million improvement in the bottom line. The need to improve and lower the fuel consumption of marine diesel engines has never been more urgent in the history of marine industry.

minimal work and near-zero down time to retrofit to the existing fuel oil system.

Emulsified fuel in simple term means a mixture of water in fuel oil in such a way that small particles of water are formed in the fuel oil to produce a stable water-in-fuel emulsion. The main purpose of doping fuel oil with small quantities of water is to improve combustion efficiency and thus achieve better specific fuel consumption of marine diesel engines. Since the invention of the diesel engine, the focus of improving fuel combustion efficiency has been to improve the mixing of fuel and air inside the combustion chamber so that a more complete combustion can take place within the time of the engine cycle. This has been achieved, among many innovative ideas, largely by better fuel injection techniques, better air flow dynamics and allowing more time for combustion through longer engine strokes etc. There is a finite limit to the improvement to the mixing of fuel and air in the combustion chamber because there are physical limitations to the fuel injector design, dynamic air flow and time available for combustion. Referring to published information from diesel engine manufacturers, for example Wärtsilä (Fig 2), the total efficiency of the best diesel engines has improved by about 6% since 1985. If we consider that slow speed diesel engines are the most efficient diesel engine and that 3% improvement came from better mechanical efficiency due to direct coupling, the improvement in combustion efficiency is about 3%. Compare this with the improvement using water-in-fuel emulsions (or emulsified fuels) which improve combustion efficiency by 3-5%, the benefits of the application of water-in-fuel emulsions are significant.

Recently, there is keen interest in LNG as the fuel of the future; it has abundant supply for the next 100 years and is clean. However, there are about 70,000 existing ships in the world that still rely on the marine diesel engine for power. Challenges of LNG application to existing ships are: the larger storage requirements (approx 4 times more) and the high cost of retrofitting. LNG is the fuel of the future for new buildings but it is unlikely to replace marine fuel oil as the primary fuel for the existing fleet.

Emulsified fuel system developed by Blue Ocean Solutions Pte Ltd (BOS), is a practical and proven technology which can improve fuel efficiency by 3-5% without any modifications to the marine diesel engines. The compelling advantage is that it is easy to retrofit to existing ships, requiring

Figure 1. bunkerworld index

1600

1500

1400

1300

1200

Feb Mar Apr May Jun Jul Aug sep oct nov dec Jan 2011 2011 2011 2011 2011 2011 2011 2011 2011 2011 2011 2012

1970 1976 1980 1985 1990 1995 2000 2006

%

50

48

46

44

42

40

development of the shaft efficiency of Wärtsilä’s best engines

Figure 2. improvement in shaft efficiency

The main benefits offered by the Blue Ocean Solutions Emulsified Fuel System (BOS EFS) are summarised in Fig 3. Other benefits include:

• Cleanerdieselenginehotcomponents (Injectors, Cylinder liners, Exhaust Valves, Exhaust Ducts, Turbochargers):

- Better efficiency

- Maintaining engine peak efficiency over longer time

- Extending time interval between overhauls which is translated into lower maintenance costs in terms of fewer spare parts and less labour

- Reduced soot blowing – save steam and less soot on open decks

- Reduced water washing – reduced amount of dirty and contaminated water that is difficult to treat prior to disposal.

UNDERSTANDING EMULSIFIED FUELIt is well known that worn out fuel injectors can cause increase in fuel consumption due to poorer atomisation i.e. larger injected fuel droplets. The key to improve combustion efficiency is by creating smaller injected fuel droplets. Smaller fuel droplets will achieve better fuel-air mixture because more fuel-air surface will be available for combustion which is the common experience shared by marine engineers. However, there are practical limitations to producing smaller and smaller fuel droplets for a required amount of fuel that must be injected to produce the required power. These are:

- Injector holes has limitation to how small they can be

- Fuel injection pressure is limited by the fuel pumps

- Fuel injection time is limited by speed of the engine cycle

A practical and proven technology which can improve fuel efficiency by 3-5% without any modifications to the marine diesel engines.

The goal of emulsified fuel is to create a secondary atomisation effect after the primary injection process, by adding water into the fuel. The secondary atomisation effect creates even smaller fuel droplets after injection.

Emulsified fuel is a homogeneous mixture of microscopic water particles-in-fuel product.Typically, the microscopic water particles in the fuel are heated in the fuel line to about 130-135 deg C under pressure of 8 – 10 bar. Referring to Table 1,

table 1. boiling point of water

Pressure (bar) boiling Point (deg c)

1.00 100

1.50 112

2.00 120

2.50 127

3.00 134

3.50 139

4.00 144

4.50 147

5.00 152


Figure 3. Main benefits of bos eFs

3% - 5%

no down time and stoppages neededeasy installation

10% - 15%*nox reduction

standard and approved parts and equipments

Fuel saving

produces emulsified fuel on demandno storage needed

reliable, low maintenance

fully automated

hfo, ifo, mdoCan handle multi fuel

ease of use

* with 10% water


All of the above methods have the following drawbacks:

• Needforfrequentmaintenance

• Rapidperformancedeteriorationduetowearof parts in the marine environment

• Unreliabilitybecauseoftheabovereasons

All the above methods also cannot control the production of optimal emulsified fuel because they continuously break the water-in-fuel particles in the re-circulated fuel to nano-sized water particles which do not create the secondary explosion effect. Too small water-in-fuel particles do not have enough energy to produce the desired explosive effect. Unlike all of the existing emulsifiers, the BOS emulsifier will reliably and consistently produce the optimal emulsified fuel needed to achieve the secondary atomisation effect. It also re-emulsifies any re-circulated water and fuel to the same optimal emulsified fuel.

BOS emulsifier was developed after extensive research, specifically for producing optimal water-in-fuel emulsions for achieving the best fuel savings. The non-dimensional emulsifier model

water will not boil at 135 deg C when the fuel line pressure is above 3.2 bar. When the water-in-fuel emulsion is injected into the engine, the superheated water particles will boil or explode instantaneously. A simplified analogy would be like the exploding effect, when water is poured on to hot oil. This is the secondary atomisation effect that creates a finer fuel mist and better fuel-air mixture for combustion inside the cylinder after the primary injection as illustrated by Fig 4.

Through extensive and comprehensive research, the optimum water content and particle sizes for maximum fuel saving on marine diesel engines were established: water content of 10% and particle sizes of 2-6µm.

KEY INNOVATION IN EMULSIFIED FUEL SYSTEMThere are several types of emulsifiers available from the market which use one of the following methods of emulsification:

• Mechanicalshearing/homogeniser

• Cavitation

• Ultra-sonicvibration

Figure 4. secondary Atomisation

Fuel droplet (approx 30 -100µm)

superheated water-in-fuel particle (2-6µm)

secondary explosion

secondary explosions create better fuel-air mixture for combusion at lower temperature

superheated microscopic water particles explode when injected into cylinder

Water-in-fuel emulsion under high pressure

(patented) was developed to ensure that BOS emulsifier can be scaled to meet the fuel flow range of any ships. The considerations of the non-dimensional model are illustrated in Fig 5. A set of design curves was empirically determined, based on the non-dimensional model. The BOS emulsifier is illustrated in Fig 6. This is an important innovation in emulsified fuel technology which ensures that the optimum water-in-fuel emulsion (Fig 7) is always produced reliably and consistently to achieve the best savings for any ships. This BOS emulsifier makes possible a practical emulsified fuel system that produces emulsified fuel on demand reliably and consistently. There is no similar emulsifier as the BOS emulsifier in the world which is protected by a world-wide patent.

Figure 7. emulsion (hFo)

the non-dimensional model is a

function of:

• mean water particle size ratio, p/D

• fuel Reynold number, (ρfVfd)/µf

• nozzle dimension ratio, d/D

• velocity ratio, Vw/Vf

• Weber number, σ/(ρfdVf2)

• Relative density, ρf/ρw

• Viscosity ratio, µf/µw

Figure 5. Considerations of bos emulsifier Model

In summary, the innovative features that differentiate BOS Emulsifier from others are:

• Producesoptimalwater-in-fuelemulsion to create the secondary atomisation effect that improves combustion

• Nomovingparts

• Designedforreliabilityandeasymaintenance

• Producesemulsifiedfuelondemandfor immediate consumption. Hence, no storage of emulsified fuel is needed. (Storage of emulsified fuel may cause biological corrosion due to growth of biological microbes)

• Re-emulsifiesre-circulatedwaterandfueltothe same optimal emulsified fuel

• Canbescaledtoproduceoptimalemulsified fuel for ships of any sizes

This is an important innovation in emulsified fuel technology which ensures that the optimum water-in-fuel emulsion is always produced reliably and consistently to achieve the best savings for any ships.


Figure 6. bos emulsifier

Water

emulsionFuel oil

➙ ➙

➙


Figure 8. eFs retrofitting layout

INSTALLATIONBOS EFS is designed for easy and quick installation. The ship does not need to dry dock. Retrofitting can be done when the ship is at port with near zero down time. Three auto change over valves as illustrated in Fig 8 and Fig 9, would be installed to the fuel line by retrofitting a small section of the FO pipe (between the circulation pumps and heaters) while the ship is at port, without interrupting the ship’s operations. The new fuel pipes can then be blanked off and the rest of the retrofitting work can be completed when the ship is at sea, if necessary. BOS EFS is modular as illustrated in Fig 10 so that installation is easy and can be quickly completed. The size of the EFS skid is approx L1.8m x W1.2m x H2.0m.

circulation pump

Booster pump

fuel flow meter

mixing tank

3 x new auto change over Valves

heaterViscometer

water

no

nc nc

diesel engine

fuel tank

note: retrofittted eps equipment are in red

efs

Figure 9. example of Fo Pipe retro-fitting

ACTUATOR VALVE (NC) DN65

THERMOMETER

PIPE, DN65

EMULSIFIER

MANUAL VALVE DN65

ACTUATOR VALVE (NC) DN65

ACTUATOR VALVE (NO) DN65

PRESSURE GAUGE

Figure 10. Modules of bos eFs


fuel

fuel flow meter

water supply

nc nc

emulsifier

3 x auto change over

Valves

emulsion

eMulsiFier Module

WAter Control Module

plc

pump + motor

water flow meter

w

ePs Controller

engine rooM

Control

WAter tAnK

Module

no

water


of a patented emulsifier which ensures that the desired water-in-fuel emulsions are produced reliably and consistently to give the best fuel savings. Additionally, BOS EFS is modular and easy to retrofit with near zero operational down-time. Dry docking is not necessary. It can be retrofitted to the existing fuel oil system without any modifications to the diesel engines. The track record of BOS spans over 20 years since the 1980’s. Fuel reductions of 3-5% have been consistently achieved without any adverse effects on the maintenance of the hot components of the diesel engine.

Figure 11. From top: Connaught; spirit of Free enterprise; neptune Crystal; and london Viscount

REFERENCES AND RESULTSThe experiences and references of BOS EFS span over 20 plus years, since the 1980’s. Dr Jerry Ng KL, the CEO and founder of Blue Ocean Solutions Pte Ltd started his research in the application of emulsified fuel technology to marine diesel engines in the 1980’s. He was the first in the world to install and test emulsified fuel systems in the 1980’s, on four ships (Fig 11):

• M/SConnaught,RoRoFerryofB+I, medium speed engine

• M/SSpiritofFreeEnterprise,RoRoFerryof Townsend Thoresen, medium speed engine

• M/SNeptuneCrystal,ContainershipofNOL, slow speed engine

• M/SLondonViscount,Bulkcarrierof London Overseas Freighters, slow speed engine

Reductions of 3-5% in fuel consumptions were achieved. The conditions of the engine’s fuel pumps, fuel injectors, cylinder liners, pistons, exhaust valves, and turbo chargers were also examined after 6 months and they were found to remain in good conditions.

In 2011, Blue Ocean Solutions Pte Ltd signed an agreement with Wartsila to provide the BOS EFS for testing under Project Hercules (http://www.hercules-b.com/1/article/english/1/2/index.htm). Results of the test based on 2-stage Turbo Charged Miller and using LFO with 10% water showed that fuel reductions of 0.5 to 4.5% were achieved. The test facility and results are shown in Fig 12.

An order for BOS EFS from a major cruise ship owner was secured in 2011. Fig 13 shows the typical fuel rate and power trend curves of the cruise ship when BOS EFS was switched on and water was injected into the fuel oil system. It can be seen that the fuel rate dropped when BOS EFS was switched on while the engine power remained constant. Fuel reductions of more than 3% were achieved.

CONCLUSION An innovative emulsified fuel system has been developed by Blue Ocean Solutions Pte Ltd to address ship owners’ concern about the high cost of fuel oil. The key innovation is in the development

Figure 13. trend Curves when eFs is on


Figure 12. Project hercules Facility and test results

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fuel: lfo Valve timing: miller

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©2012 Keppel Offshore & Marine Technology Centre Pte Ltd (“KOMtech”) Cover Image: A 3D rendering of a Small-Scale LNG distribution in operation - LNG transfered by KOMtech’s Cryogenics Hose Handling system between an LNG Carrier and a floating LNG Terminal

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31 Shipyard Road

Singapore 628130

Tel: (65) 6591 5450

Fax: (65) 6265 9513

Email: [email protected]

Co Reg No: 200615559N

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